Multiple Automated Reactor Systems (MARS). 2. Effect of Microreactor

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Multiple Automated Reactor Systems (MARS). 2. Effect of Microreactor Configurations on Homogeneous Gas-Phase and Wall-Catalyzed Reactions for 1,3-Butadiene Oxidation Patrick L. Mills* DuPont Company, Experimental Station, E304/A204, Wilmington, Delaware 19880-0304

Jacques F. Nicole Catalytica Energy Systems, 430 Ferguson Drive, Mountain View, California 94043-5272

Wall-catalyzed and homogeneous gas-phase reactions for the gas-phase oxidation of 1,3-butadiene are studied using the Multiple Automated Reactor System (MARS) described in part 1 of this series of papers to assess reaction inertness for a typical parallel catalytic reactor system. The heterogeneous catalyzed version of 1,3-butadiene oxidation represents an alternate C4 hydrocarbonbased route for the manufacture of a tetrahydrofuran (THF) monomer. Polyether glycol polymers derived from THF are used to manufacture various commercial materials where elastic or flexible impact-resistant properties are desirable. Experiments were performed using parallel banks of the MARS U-tube and straight-through-tube reactor designs in both empty tube and packedbed modes. Empty tube mode data using typical heterogeneous catalyst reaction conditions showed the presence of significant reactant conversion to combustion products for both reactor designs. The U-tube design generated different reactant conversion results when compared to the straight-through-tube design. These differences were attributed to the interaction between system design features and reaction conditions, such as the reactor tube geometry, the product transfer line design, and the contact times in various heated zones. Various materials of construction and catalyst bed retaining materials were also evaluated for inertness using the packed-bed mode. When a butadiene-rich feed gas was used, the O2 conversion was generally higher in the presence of various metals versus glass according to type 316 stainless steel > Ti > glass wool > glass beads. The results suggest that caution must be exercised when catalyst performance data are measured using parallel gas-phase reactor systems, especially for substrates with functional groups that may react in the gas phase and in the presence of various materials of construction. Each reactor unit in the parallel bank should be evaluated for reaction inertness before a catalyst development campaign is initiated to avoid incorrect conclusions on catalyst ranking and generation of biased data that contain false measures of catalyst performance. 1. Introduction Tetrahydrofuran (THF) is a key monomer used in the manufacture of poly(tetramethylene) ethylene glycol (PTMEG) polymers. PTMEG polymers are key ingredients used to manufacture Spandex® fibers and other industrial applications where flexible impact-resistant properties are required.1 A recent process economics analysis compared various routes for the manufacture of THF to develop an improved understanding of the challenges associated with the use of traditional hydrocarbons versus biomass-derived materials as feedstocks.2 One key conclusion was that, among various multiple-step routes for THF manufacture based on C3 and C4 hydrocarbons, the gas-phase partial oxidation of 1,3-butadiene with furan as an intermediate provided a competitive alternative to processes based on n-butane oxidation where maleic anhydride (MAN) is the intermediate. A comparison between the main reaction steps for these two competing routes is shown in Figure 1. * To whom correspondence should be addressed. Tel.: (302) 695-8100. Fax: (302) 695-3501. E-mail: Patrick.L.Mills@ usa.dupont.com.

Figure 1. Gas-phase oxidation routes to THF using either 1,3butadiene or n-butane as the feedstock.

Table 1 compares the key process parameters for the two routes, which reveals some notable differences. The data presented here assume that both routes are practiced using a circulating solids reactor that is equipped with separate riser-regenerator zones similar to the first reaction stage of the DuPont THF process.3 In the riser zone of this process, the C4 hydrocarbonrich feed gas is contacted with an attrition-resistant form of the metal oxide catalyst,4 where it is partly converted to the either furan or MAN via lattice oxygen through the Mars-van Krevelen mechanism.5 The metal oxide catalyst undergoes a net reduction in the oxidation state, which creates oxygen vacancies in the solid-state surface and subsurface lattices. These vacancies in the metal oxide lattice are then replenished by

10.1021/ie048830z CCC: $30.25 © 2005 American Chemical Society Published on Web 07/12/2005

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Table 1. Comparison between 1,3-Butadiene and n-Butane Processes to THF process route parameter O2 stoichiometry, mol catalyst oxocapacity H2 stoichiometry, mol pressure, psi furan selectivity, % MAN selectivity,a % C4 conversion,a % furan yield, % MAN yield, %

1,3-butadiene to THF

n-butane to THF

1 molybdate (NaBiAsMoO) 10 times 2 2000 70b 40b 28

a

Based on data extracted from refs 1-3. b The 1,3-butadiene conversion and furan selectivity correspond to those of the catalyst from patent example 5 in ref 4.

contacting the catalyst with air in a bubbling-bed regenerator zone before the attrition-resistant catalyst is returned to the riser zone. Additional details on the role of surface chemisorption and the catalyst lattice in the reaction mechanism for the partial oxidation of n-butane to MAN over vanadium-phosphorus oxide (VPO) have been recently reported.6 The indicated conversion-selectivity parameters for the partial oxidation of 1,3-butadiene are based upon data for a Na2.1Bi6.3AsMo9O40 mixed oxide catalyst,7 while the corresponding parameters for n-butane oxidation were obtained from the open literature3 and patent examples4,8 for a VPO catalyst. The data in Table 1 suggest that the advantages of using 1,3-butadiene versus n-butane include reduced consumption of gasphase oxygen per mole of C4 hydrocarbon reactant (1 mol vs 3.5 mol), greater lattice oxygen capacity of the bismuth-molydenum oxide catalyst (10 times vs 1 times), reduced consumption of hydrogen per mole of THF (2 mol vs 5 mol), and a reduced hydrogen pressure for conversion of furan versus maleic acid to THF (2000 psi). An important implication of the reduced consumption of gas-phase oxygen and a higher oxocapacity is that catalyst regeneration is simplified when either more modern fluidized-bed or circulatingsolids reactor systems are utilized.3 However, the Na2.1Bi6.3AsMo9O40 catalyst for the 1,3-butadiene oxidation process produces a lower yield to furan (5.8%)7 when compared to the overall yield of MAN (28%) from the n-butane route.3 Discovery of an improved 1,3-butadiene oxidation catalyst with a furan yield that competes with those obtained from existing catalysts for n-butane to MAN would create a competitive advantage for this route. Catalyst screening for the partial oxidation of 1,3butadiene to furan and MAN is more complicated than some other partial oxidation reactions because homogeneous gas-phase and wall-catalyzed reactions can disguise the true heterogeneous catalyst performance parameters. This is exemplified by early work on catalyst discovery using 1,3-butadiene and 1-butene as the C4 hydrocarbon feedstocks in laboratory-scale fixed beds, which resulted in the identification of modified bismuth-molybdenum oxides as the preferred catalysts for furan.7 Experiments using feed mixtures containing 1,3-butadiene, 1-butene, and O2 at temperatures between 500 and 550 °C showed that U-tube laboratoryscale reactors constructed from either gold or type 316 stainless steel (SS) resulted in undesired wall-catalyzed

reactions, whereas those fabricated from titanium, aluminum, and quartz were nearly inert. For example, when a feed stream containing 25% 1,3-butadiene, 3.5% 1-butene, 16% O2, 10% H2O, and a balance (45.5%) of helium was introduced through a 1/4-in.-o.d. titanium U-tube reactor at 525 °C using a flow rate of 425 mL‚min-1, the total conversion of the hydrocarbons was 0.7%. Conversely, when a 1/4-in.-o.d. type 316 SS U-tube was substituted for the titanium tube, the hydrocarbon conversion increased to ca. 5% with 65-80% of the product gas consisting of CO2 even though the temperature was lower by 11 °C (i.e., 514 °C). Additional performance data showed that titanium was relatively inert to the reaction mixtures, so it was the preferred material of construction for this particular C4 hydrocarbon partial oxidation over the range of process conditions that were studied. However, these experiments were conducted using a traditional single-tube fixed bed whose reactor dimensions, reactor metal quality, and other reactor system design parameters and operating conditions differed from those used in the current study. For this reason, a relevant comparison is not possible. The effect of different materials of construction on the partial oxidation of light olefins and other substrates in tubular reactors was investigated by Albright and coworkers in the late 1970s.9-11 Oxidation reactions using methane, propane, propylene, CO, oxygenated compounds, and hydrogen as substrates were performed using laboratory-scale reactors constructed of aluminum, copper, Pyrex, and carbon steel over a temperature range of 300-500 °C and at atmospheric pressure. Both new and aged surfaces were investigated. The rate of oxidation and decomposition of oxygenated compounds varied according to the sequence of copper > steel > aluminum > Pyrex, where copper exhibited the greatest activity. The rate of decomposition was also enhanced by aged or used metal surfaces when compared to a new surface. It was also shown that the addition of water vapor to the feed gas reduced the number of surface sites that promoted the oxidation and decomposition reactions. However, when CO was also present with water vapor, both hydrogen and CO2 were produced, which indicated that the water gas shift reaction had occurred. Figure 3 in one of these works9 shows that, at 360 °C, an oxygen conversion of 5% is obtained in a new steel reactor versus 85% in an aged reactor when a propaneoxygen feed gas molar ratio of 5:1 and a residence time of 2 min were used. The tubular reactor was constructed of low carbon steel with an i.d. of 0.62-0.77 cm, a length of 244 cm, and an internal volume between 74 and 114 cm3. The resulting surface-to-volume ratio was 5.2-6.4 cm-1. The geometries of both of the MARS-2 and MARS-3 reactor tubes described in part 1 of this series of papers (dr ) 5.35 mm) have a surface-to-volume ratio of 7.5 cm-1, which exceeds those used in the above-cited work of Albright and co-workers. This suggests that wall-catalyzed reactions could be important for the MARS-2 and MARS-3 reactor systems when used for hydrocarbon oxidation reactions so that an assessment of their contribution merits investigation. From a broader perspective, a key requirement for obtaining reliable data for gas-solid heterogeneous catalyzed reactions from laboratory-scale reactors is that the rates of homogeneous gas-phase and wall-catalyzed reactions should be negligible when compared to those for the solid-catalyzed reactions.12 Gas-phase hydrocar-

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bon oxidation reactions are particularly susceptible to undesired side reactions in laboratory-scale reactors because of the existence of various reaction channels.5 These nonselective reactions can generally occur at two primary reactor system locations: (1) in the gas bulk of unpacked heated zones that are present near the feed gas inlet and product gas exit of the catalytic bed; (2) on the walls of the heated reactor and gas transport tubing. In the latter case, the materials of construction used to fabricate these components may contain surface heterogeneity and be comprised of various metals that collectively serve as catalytic sites for wall-catalyzed destructive reactions. Most open literature studies on the catalyst science for gas-phase oxidation chemistry do not include an assessment of the relative contribution of homogeneous gas-phase and wall-catalyzed reactions to the observed solid-catalyzed reactions. This is one possible explanation why different conclusions on the catalyst performance are sometimes reached when the same catalyst composition is evaluated using different reactor systems. Typical materials of construction used for fabrication of reactors and other heated components in reactor systems, such as gas transport tubing and gas-sampling systems, can have particular importance when highthroughput methods are utilized for catalyst screening.13 These methods employ reactors that use smaller catalyst charges than their larger single-reactor counterparts. This trend has pushed reactor diameters from conventional sizes (e.g., 5-6 mm) to smaller dimensions (e.g., 0.5-1 mm or less), which translates into larger ratios for the reactor tube surface area to reactor volume (e.g., 10 vs 1). Consequently, these smaller diameter reactors increase the potential contribution of wallcatalyzed reactions relative to those in the gas-phase bulk when compared to conventional larger diameter laboratory-scale reactors. In addition, steel tubing and other heated components used in the fluid transport subsystems of high-throughput systems also have smaller inner diameters (e.g., 0.762 mm i.d.) when compared to those used in more conventional systems (e.g., 4.572 mm i.d.). This feature also generates higher specific surface areas (6 vs 1), which again increases the potential contribution of wall-catalyzed reactions. However, an assessment of the relative importance of undesired wallcatalyzed and bulk gas-phase side reactions relative to the desired catalytic oxidation reaction is not always reported in the open literature, especially in more recent publications concerned with applications of highthroughput testing for partial oxidation catalysts. The importance of identifying an inert reactor material in high-throughput testing for partial oxidation catalysis was emphasized in a recent review,14 although no specific data were provided. In part 1 of this series of papers, a reactor system called the Multiple Automated Reactor System (MARS) was described for detailed evaluations of gas-phase partial oxidation systems. The system contains a parallel bank of six independently controlled fixed-bed microreactors. The MARS-2 prototype reactor design was based upon a metal U-tube reactor configuration whose temperature was controlled using a fluidized-bed sand bath. A replicate of the same system called MARS-3 was also described that is similar to the MARS-2 prototype in nearly all aspects except for the reactor tube configuration and the hardware used for independent control of the reactor tube temperature. The MARS-3 replicate

also incorporated a metal reactor, but the configuration consisted of a straight-through tube whose temperature was controlled by a custom-designed split furnace. Both the MARS-2 prototype and MARS-3 replicate systems were intended to provide the same performance data when the same catalyst array was independently tested using an identical set of reaction conditions and operating protocol. However, in view of the work of Albright and co-workers,9-11 these two different reactor designs could potentially generate different performance results for a given partial oxidation reaction with the same catalyst even though all other supporting system components and operating protocols are identical. Identification of alternate routes to the THF monomer2 was one motivating factor for the current study. Catalyst screening efforts for the partial oxidation of 1,3-butadiene to furan using both the MARS-2 and MARS-3 systems showed that different performance results were generated when the same catalyst compositions were evaluated under identical reaction conditions. Because the MARS-2 and MARS-3 reactor systems differed mainly in the reactor tube configuration and length of the product gas transport tubing, the presence of homogeneous gas-phase and wall-catalyzed reaction oxidation was suspected as a possible reason for the differences in the catalyst performance. When MARS-2 and MARS-3 were used for the invention of novel catalysts for n-butane oxidation to MAN,15-17 similar performance data were obtained from the two independent parallel reactor systems. Experiments conducted with empty reactor tubes, or with tubes filled with various materials of construction, showed that no measurable reactions occurred when n-butane and oxygen mixtures were contacted using the same conditions that were employed in catalyst screening work. This suggested that the observed catalyst performance was solely due to the solid-catalyzed reactions, which differed from the case for 1,3-butadiene oxidation. The primary objective of this paper is to compare the relative contributions of both homogeneous and wallcatalyzed reactions in two independent laboratory-scale parallel reactor systems that are used for detailed performance evaluations of gas-phase partial oxidation heterogeneous catalysts. While the presence of these phenomena was demonstrated in previous work using larger-scale, single-tube reactors with a few substrates, they have not been quantified for any commercially relevant reactions using more modern approaches for catalyst development based on smaller-scale, parallel reactor systems. Various designs for parallel catalytic reactor systems have emerged within the past decade and new versions continue to appear13,14,18 so that catalytic chemistry is being investigated at a higher rate and with a greater degree of sophistication than ever before possible. Consequently, the elimination of unwanted reactor system effects is especially critical to maintain data quality and to ensure that catalyst science, reactor design, and catalytic process development proceed on a sound basis. Two different MARS fixed-bed microreactor system designs were used to study the oxidation of 1,3-butadiene over a range of contact times and reaction temperatures as part of an effort to assess the role of wallcatalyzed and gas-phase homogeneous reactions. Both the MARS U-tube and straight-through-tube reactor designs either were operated as empty tubes or were filled with various materials of construction in granular

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form. The reactivity of materials that are commonly used to hold catalyst beds in place, such as glass wool and type 316 SS sintered porous disks, was also studied. The effect of excess steam on reactivity was also briefly investigated. Finally, an attempt was also made to quantify the observed rates of reaction for both oxygen and 1,3-butadiene under a variety of reaction conditions using data from both empty reactor tubes and tubes filled with various materials of construction. It is shown that the classical U-tube reactor design generates a greater yield of undesired byproducts when compared to a conventional straight-through-tube reactor design for this particular application. A detailed study of the oxidation kinetics for 1,3butadiene lies outside the scope of the present work, especially because an extensive literature already exists in connection with the mechanism of hydrocarbon combustion and comprehensive models have been recently developed.19 The global kinetics extracted here are intended to provide a common basis for comparing the performance data collected from the two independent reactor configurations. This will be a useful starting basis when evaluating the next-generation parallel microreactors where undesired reactions are mitigated. 2. Experimental Section 2.1. Reactor Systems. All experiments were performed using either the MARS-2 or MARS-3 reactor systems. The detailed descriptions of these systems and other operating features were provided in part 1 of this series of papers. In some cases, only a subset of the six reaction tubes in the systems were used because these were indicative of the data obtained with the remaining tubes. 2.2. Materials. To assess the possible contribution of wall-catalyzed reactions and generate additional surface area, various materials of construction were loaded into the reactor tubes in the form of irregularshaped chips. The bulk materials were purchased from commercial sources and included type 316 SS, titanium, and quartz glass. The metals chips were produced by sawing the block material into various slices using a coarse-toothed blade. The resulting chipped material was then collected and sieved into various size fractions. Before being loaded into the reactor, the material was washed several times with acetone to remove any substances that could act as possible contaminants. 1,3-Butadiene was purchased from a commercial source (Roberts Oxygen Co., West Chester, PA) as a blended mixture containing 10% 1,3-butadiene, 10% N2, and 80% (balance) He. Oxygen was also purchased as a blend whose composition was 10% O2, 10% N2, and 80% (balance) He. The choice of N2 allowed this component to be used as an internal standard for any components that were detected by the thermal conductivity detector. Additional details on the analytical method and the techniques used for translation of the raw gas chromatography (GC) data into various measures of composition have been previously described. 20 3. Data Analysis The reactor performance data given below were interpreted using the measured values for the overall conversions of oxygen and 1,3-butadiene. The intention here was to provide insight into the relative contribution of the selective versus nonselective reactions and to assess which reactions might be active in the gas bulk

Table 2. Stoichiometric and Species Matrices for the Cases Where 1,3-Butadiene Is Oxidized to Furan, CO, and CO2, Furan to CO and CO2, and CO to CO2 elements of the stoichiometric matrix, Rji 1 0 0 -1 -1 0

2 3 3 2 1 0

0 4 0 4 0 1

0 0 4 0 4 -1

-1 -1 -1 0 0 0

-1 -5.5 -3.5 -3.5 -2 -0.5

species, Ai C4H4O (furan) H2O CO2 CO C4H6 O2

and on the exposed metal surfaces. This straightforward approach is particularly useful in the analysis of oxidation reaction kinetics because it allows an overall assessment of O2 utilization. A more detailed approach could be implemented that involves the use of reaction extents, which would allow O2 utilization to be determined for each independent reaction. This would require information such as kinetic rate expressions for the homogeneous gas-phase and heterogeneous reactions, volumes for the various empty tube zones, the void volume for the packed-bed zone, temperature profiles in the various heated zones, and gas contact times, to name a few parameters. The results of this analysis would have produced more detailed information on the quantification of the undesired reaction channels, which could be useful in possible future work involving the design of inert reactor systems. However, the conclusions regarding the importance of identifying the presence and magnitude of the undesired reactions would be unchanged. For this reason, a less complicated approach was followed as outlined below. For a single reaction having the stoichiometry

aA + bB f P

(1)

it can be shown that the conversion of A is related to the conversion of B according to

XA )

a FB0 X b FA0 B

(2)

where FA0 and FB0 are the molar flow rates of species A and B at the reactor inlet, respectively. The case where multiple reactions are involved is more complicated and requires the use of the species vector and stoichiometric matrix to describe the reaction network. Table 2 shows the members of the stoichiometric matrix R j T and the species vector A h for the set of chemical equations that satisfy R j TA h ) 0 for the case where 1,3-butadiene is oxidized to furan, CO, and CO2. The combustion reactions of furan to CO, furan to CO2, and CO to CO2 are also included, but only three reactions (bolded) are independent because the other three can be formed by a linear combination. If the partial oxidation of 1,3-butadiene to furan was the only reaction occurring, then eq 2 shows that a plot of oxygen conversion (XA) versus the 1,3-butadiene conversion (XB) would yield a straight line whose slope is (a/b)(FB0/FA0), where R15 ) R16 ) -1 as shown in Table 2. All other reactions between oxygen (species A) and 1,3-butadiene (species B) have a stoichiometric coefficient ratio of a/b that is greater than unity. For example, the combustion of 1,3-butadiene to CO would have a stoichiometric coefficient ratio of 3.5/1 ) 3.5, while the combustion of 1,3-butadiene to CO2 would have a stoichiometric coefficient ratio of 5.5/1 ) 5.5. The results in a later section


indicate that partial and total combustion reactions are dominant because the stoichiometric coefficient ratios are primarily in the range of 3-5. 4. Results and Discussion This section describes the key findings obtained when a feed stream containing 1,3-butadiene and oxygen is introduced to the two different MARS reactor designs over the same range of temperatures and contact times that were used in catalyst screening experiments. The first section compares the performance using both empty tubes and tubes filled with various materials of construction. The next section examines the catalytic activity of porous metal disks and glass wool used to maintain the catalyst bed in place. The third section provides estimates of kinetic rate parameters for the combined homogeneous and wall-catalyzed reactions. The final section quantifies the reactor temperature profiles and provides data that show how these are affected by the gas flow rates or contact time and steam addition. 4.1. Effect of the Reactor Tube Configuration. A direct comparison between both the MARS-2 U-tube and MARS-3 straight-through-tube reactor designs was performed either by operating these units in the empty reactor tube mode or by filling each reactor tube with a known quantity of either glass chips, titanium, or type 316 SS. In all cases, the experiments were performed using the same contact time (gas flow rate), feed composition, and temperature. The key results are summarized below. 4.1.1. Comparison of Reactor Volumes and Metal Surface Areas. Reactor volumes and exposed metal surface areas for the MARS-2 U-tube and MARS-3 straight-through-tube reactor designs are compared to gain insight into the magnitude of the differences between their design parameters. The empty reactor volume will impact the magnitude of the homogeneous gas-phase reactions, while the total exposed metal surface area will contribute to heterogeneous wallcatalyzed reactions. The exposed metal surface area includes the reactor wall area as well as the surface area of any granular material that is added to the empty reactor tube. The latter approach provides a method for assessing the catalytic activity of a given material relative to that obtained from open empty tube operation. The results given below will show that the trends in the oxygen and 1,3-butadiene conversions obtained from each reactor can be correlated to the reactor physical parameters and thermal characteristics, such as the overall tube length, length of the heated zones, and temperature profiles in the heated zones. To assess the relative impact of the metal surface area on the heterogeneous reactions, it is useful to examine the ratio of the metal wall surface area that is external to the catalyst zone to the total metal surface area in the catalyst bed zone. The total metal surface area in the catalyst bed zone includes the wall surface area (i.e., the reactor wall surface area that is confining the catalyst) and the external surface area of the metal packing or catalyst. This ratio is defined by the following relationship:

R)

Sw ) Sp

LE/LC 6 w 1 1+ dp Fp π dILC

(3)

Figure 2. Graphical representation of eq 3 showing the effect of the length of the external catalyst zone on the ratio of the wallto-particle surface area. The points correspond to the effective lengths of the MARS-2 U-tube and MARS-3 straight-through-tube reactors. Table 3. Metal Surface Area Ratios in Empty versus Chip-Filled Reactorsa reactor mode

w [g]

empty type 316 SS titanium

0 0.88 0.5

Fp [g/mL]

R-MARS-2

R-MARS-3

8.02 4.51

1 ∼6 ∼6

1 2.79 2.77

a MARS-2 reactor dimensions: d ) 0.47 cm; L ) 2 cm; L = I C E 40 cm; dp ) 1 mm assumed. MARS-3 reactor dimensions: dI ) 0.47 cm; LC ) 2 cm; LE = 18 cm; dp ) 1 mm assumed.

The notation used in eq 3 is as follows: Sw ) wall surface area, Sp ) particle external surface area, dp ) mean diameter of the catalyst or packing based on the equivalent volume of a spherical particle, w ) catalyst or packing weight, Fp ) catalyst or metal packing density, dI ) reactor inner diameter, LC ) length of the catalyst bed zone, and LE ) length of the tubing external to the catalyst bed zone. The latter parameter can exceed the length of the entire reactor tubing assembly (i.e., the inlet and exit sections of the tubing on either side of the catalyst bed zone and the catalyst bed zone itself), depending upon the temperature profiles that exist on both sides of the catalyst bed zone. The above equation is plotted in Figure 2 using the length of the tubing external to the catalyst bed zone as a parameter, with fixed values for the remaining variables. The particular fixed parameters used here are shown in Table 3 for both the MARS-2 and MARS-3 reactors, which includes typical charge weights for the metal granules. The external reactor tube length for the MARS-2 U-tube design (40 cm) is about a factor of 2 greater than the corresponding tube length for the MARS-3 straight-through-tube design(18 cm). These values are indicated on the figure by the points. When the reactors are filled with the indicated amounts of metal granules, the wall-to-particle surface area ratio (R) in the MARS-2 U-tube reactor is greater than the corresponding one for the MARS-3 reactor by the same factor because the external tube length appears as a linear parameter in eq 3. Greater values for the reactor volume and exposed metal surface area in the MARS-2 U-tube reactor would generate higher rates for the homogeneous and wall-catalyzed reactions when compared to the MARS-3 straight-through-tube reactor. The oxygen and 1,3-butadiene conversion data given in a later section for the two different reactor designs support this assertion. 4.1.2. Empty Tube Operating Mode. Figure 3 compares the oxygen conversion (Figure 3a) and the 1,3butadiene conversion (Figure 3b) versus temperature

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Figure 3. Effect of the temperature on conversion in empty MARS-2 and MARS-3 reactors: (a) effect of the temperature on oxygen conversion; (b) effect of the temperature on 1,3-butadiene conversion. Feed composition: 8.7% C4H6, 9.7% O2, 9.1% N2, balance of He. Flow rate ) 15 mL/min (NTP).

data obtained when the MARS-2 and MARS-3 reactors were operated in the empty tube mode. The MARS-2 data correspond to reactors 5 and 6 (data labeled R5 and R6), while the MARS-3 data correspond to reactors 4-6 (data labeled R4, R5, and R6). The conversion data at a given temperature for the MARS-2 U-tube reactors are always greater than the corresponding data for the MARS-3 straight-through-through reactors. For example, the oxygen conversion from the MARS-2 reactors approaches 100% at a reaction temperature of 450 °C, whereas the average oxygen conversion from the MARS-3 reactors is less than half at 40%. Because the MARS-2 U-tube reactor is immersed in the fluidized-bed sand bath, the reaction mixture is exposed to a longer heated zone when compared to the MARS-3 straight-throughtube reactor design (40 vs 18 cm) so that a higher overall conversion would be expected. Another noteworthy aspect is the magnitude of the relative differences that exist between the reactant conversion data for the MARS-2 and MARS-3 reactors. In principle, the conversion and selectivity data at a given temperature for a given reactor configuration should be identical. Figure 3a shows that the oxygen conversion data for the MARS-3 reactors (reactor tubes 4-6) slightly increase at a given temperature. The corresponding conversion data for the MARS-2 U-tube reactors (reactor tubes 5 and 6) at a given temperature exhibit a greater absolute difference. The relative differences in conversion can be qualitatively explained by the lengths of the heated transfer tubing used for transport of the product gas from each of the MARS reactors to the oven that houses the gas sampling valves. The MARS-2 reactors are arranged in a planar staggered layout relative to the gas sampling valve oven

(see Figure 2 in part 1 of this series of papers). This arrangement results in a heated transfer line for reactor 5 that is about a factor of 1.5 times greater than the corresponding one used for reactor 6. The heated transfer line behaves as an empty tubular reactor and contributes to the homogeneous and wall-catalyzed reactions. The longer section of the heated transfer line for reactor 5 would be expected to exhibit a greater relative contribution to the overall conversion when compared to the shorter section associated with reactor 6. Conversely, the MARS-3 reactors are arranged in a vertical symmetrical pattern (see Figure 3 in part 1 of this series of papers) with heated transfer lines whose lengths from reactors 4-6 to the valve oven increase by only about 10% from one reactor to the next. Consequently, the impact of the increasing length of the heated transfer tubing for the MARS-3 system on the overall conversion would be less when compared to that used in the MARS-2 system. The data in Figure 3 correlate with this assertion. The above data for the MARS-2 U-tube and MARS-3 straight-through-tube designs provide some useful insights into reactor configurations and transfer line design for 1,3-butadiene oxidation catalyst testing and other chemistries that may undergo homogeneous gasphase and wall-catalyzed reactions. Obviously, the zone containing the catalyst should be maintained as isothermal or nearly isothermal within a few degrees of the desired set point. Either direct measurements of the internal catalyst bed temperature gradients using microthermocouples (e.g., 0.2 to 0.3 mm i.d.) or local estimates of the temperature gradients using standard relationships23,25 can be used. To minimize or eliminate homogeneous and wall-catalyzed reactions in the external catalyst zone reactions, the inlet and exit zones on either side of the catalyst bed should be maintained at the minimum temperature required to maintain the reactants and products in a vapor-phase state. If THF and water were the heaviest-boiling components and they were present at dilute concentrations, the temperature of these zones could be on the order of 80-100 °C so that wall-catalyzed and homogeneous gas-phase reactions would be inoperative. However, a complicating factor is that THF can be catalytically oxidized in series to form MAN. The formation of MAN requires a notably higher transfer line temperature to eliminate vaporphase condensation, such as 175-200 °C. The elimination of heated transfer lines by a direct fluidic connection to the GC analytical system is desired, but this would require special-purpose gas-sampling hardware. This is a topic for future consideration, but it points to the complexities that can arise for a given catalyst development application. 4.1.3. Empty Tube versus Filled Tube Operating Mode. a. MARS-2 U-Tube. Data that provide some insight into the role of homogeneous gas-phase and heterogeneous metal-catalyzed reactions for the MARS-2 U-tube reactor setup are shown in Figure 4. Here, the oxygen conversion versus temperature data (Figure 4a) and 1,3-butadiene conversion versus temperature data (Figure 4b) for an empty MARS-2 U-tube reactor configuration are compared to data obtained when the reactor is filled with either type 316 SS or titanium chips. For both the empty tube and the tubes filled with the metal chips, oxygen gas, which is the limiting reactant, approaches 100% conversion as T f 460 °C, while the butadiene conversion is ca. 20%. The differ-


Figure 4. Comparison of conversion versus reaction temperature data when the MARS-2 reactor is operated as an empty tube or filled with either type 316 SS or titanium chips: (a) oxygen conversion versus reaction temperature; (b) butadiene conversion versus reaction temperature. Feed composition: 8.7% C4H6, 9.6% O2, 9.1% N2, balance of He. Flow rate ) 24 mL/min (NTP).

ences between the reactant conversions for the empty tube versus those containing metal chips depend on the temperature range. The titanium chips appear to have the greatest impact on the activity because this material produces the highest conversion over the entire temperature range. The SS chips generally produce the lowest activity over the lower range of temperatures. However, the activity imparted by this material rapidly increases as T f 460 °C, where the oxygen approaches complete conversion. The oxygen conversion from the empty tube appears to follow the data for SS over the lower temperature range, but it then undergoes a rapid increase at ca. 435 °C, where it approaches the titanium data. A key conclusion from the above data is that the MARS-2 U-tube design is not inert over the indicated range of temperatures for this application, which also corresponds to the range of interest for catalyst testing. Even if the maximum reaction temperature is restricted to 400 °C, the oxygen conversion exceeds 20% in all operating modes. Use of an alternate material in lieu of type 316 SS for the reactor tube, such as titanium, is not recommended either because this material also imparts additional catalytic activity to the butadieneoxygen reaction mixture. According to the data shown above in Table 3, the MARS-2 U-tube reactor design has the greatest reactor wall surface area when compared to the MARS-3 design. As a result, the wall-catalyzed reactions generated from the MARS-2 reactor would represent a greater relative contribution to the overall oxygen and butadiene conversions versus those provided by the MARS-3 straight-through-tube design. In addition, the contribution of the heated transfer lines cannot be neglected, as suggested previously by the results in Figure 3. b. MARS-3 Straight Tube. Data that compare the role of homogeneous gas-phase and heterogeneous

Figure 5. Comparison of conversion versus reaction temperature in an empty reactor to that of a reactor filled with glass beads, type 316 SS chips, or titanium chips: (a) oxygen conversion versus reaction temperature; (b) butadiene conversion versus reaction temperature. Feed composition: 8.7% C4H6, 9.7% O2, 9.1% N2, balance of He. Flow rate ) 30 mL/min (NTP). MARS-3 reactor 4.

metal-catalyzed reactions in the MARS-3 straightthrough-tube reactor are shown in Figure 5. Here, oxygen conversion (Figure 5a) and butadiene conversion (Figure 5b) versus temperature data for an empty tube are compared to data for tubes packed with either glass, titanium, and type 316 SS chips. The experimental conditions are summarized in the figure caption. The MARS-3 empty reactor tube data clearly show that the reactor environment is not inert under the indicated conditions once the temperature exceeds ca. 350 °C. Figure 5a shows that the oxygen conversion exceeds 80% as T f 460 °C, while Figure 5b shows that the butadiene conversion is ca. 20%. For the butadiene feed composition indicated above, oxygen is the limiting reactant so complete conversion of oxygen is possible. The MARS-3 results generally differ from those given in Figure 4 for the MARS-2 data because the MARS-3 data are clustered into two groups whereas the MARS-2 data are generally clustered together. It is also evident that the MARS-3 reactor tube, when filled with type 316 SS granules, exhibits significant reactivity when compared to the other materials of construction once the temperature exceeds ca. 360 °C. At 440 °C, the oxygen conversion (see Figure 5a) increases from 10% to 66% in the following order of the materials of construction: glass beads < titanium < empty reactor < type 316 SS. The trends for the 1,3-butadiene conversion (see Figure 5b) are the same as those for oxygen, although the overall relative increase in the 1,3-butadiene conversion is less than the overall relative increase in the oxygen conversion because butadiene is not the limiting reactant.

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Figure 6. Oxygen conversion versus 1,3-butadiene conversion in an empty reactor and a reactor filled with type 316 SS or titanium chips. Conditions: butadiene-rich feed gas containing 8.7% C4H6, 9.6% O2, 9.1% N2, and balance (72.6%) of helium; 1,3-butadiene/ oxygen feed ratio (FB0/FA0) ) 0.90; flow rates (Qg) ) 15, 30, and 60 mL/min (NTP); T ) 320-480 °C; reactor ) MARS-3 reactor 4.

The above results demonstrate that the MARS-3 straight-through-tube design is relatively inert up to ca. 400 °C because the maximum oxygen conversion does not exceed 5%. This is a notable difference from the MARS-2 U-tube design because the oxygen conversion for the latter (see Figure 4a) exceeded 20%. To safely neglect the effect of homogeneous and wall-catalyzed reactions during catalyst testing at this upper temperature limit, the catalyst activity should be high enough so that the oxygen conversion can exceed 50% by the proper selection of the contact time, catalyst loading, and other related parameters. The results in Figure 5 for the MARS-3 reactor design can be used to obtain insight into which reactions in the oxidation network in Table 2 are operative. This can be performed by first plotting the oxygen conversion versus butadiene conversion data so the stoichiometric coefficient ratio can be extracted in accordance with eq 2. Figure 6 shows the results from this analysis. The solid lines were obtained by linear regression of the butadiene conversion-oxygen conversion data. The indicated apparent stoichiometric coefficient ratios (a/ b)app were extracted from the regression analysis slopes. The ratio of the apparent oxygen to butadiene stoichiometric coefficients (a/b)app for type 316 SS is 5.06 (see eq 1). This approaches the theoretical value of (a/b)app ) 5.5 corresponding to the complete combustion of butadiene to CO2. Similarly, the value for (a/b)app for the empty reactor is 3.59, which approaches the theoretical value of (a/b)app ) 3.5 corresponding to the complete combustion of butadiene to CO. When the reactor was packed with titanium chips, the value for (a/b)app is 3.55, which approaches the value obtained for an empty reactor. The differences in conversion between the empty reactor and a reactor filled with either titanium or type 316 SS chips are small and could be attributed to deviations in the experimental variables (e.g., flow rate, temperature, GC analysis, etc.) and contact time. It can be concluded that while a reduction in the consumption of oxygen occurs using titanium versus type 316 SS, the MARS-3 reactor system is not inert over the entire range of temperature. However, as suggested above, it could be used for catalyst testing with small errors at temperatures up to ca. 400 °C under integral conditions. The corresponding product distribution ratios of CO to CO2 that were generated from the partial and total

Figure 7. Comparison of product selectivity ratios versus reaction temperature in an empty reactor to that of a reactor filled with glass beads, type 316 SS chips, or titanium chips: (a) CO/CO2 selectivity ratio versus reaction temperature. (b) Selectivity ratios between partial oxidation products (furan, MAN, and acetone) and combustion products (CO + CO2) versus reaction temperature. Feed composition: 8.7% C4H6, 9.7% O2, 9.1% N2, balance of He. Flow rate ) 30 mL/min (NTP). MARS-3 reactor 4.

combustion of 1,3-butadiene in the MARS-3 reactor are shown in Figure 7a. The relative ratio of CO to CO2 shows that type 316 SS has the lowest ratio because the production of CO2 is greatest, which was confirmed earlier in terms of the apparent stoichiometric coefficient ratio (a/b)app. The presence of a relative minimum and maximum in the product distributions at intermediate temperatures was duplicated in separate experiments and is suggestive of competing reaction pathways. Figure 7b shows how the ratio of the sum of the selectivities of the partial oxidation products, namely, furan, MAN, and acetone, to the sum of the selectivities of the COx’s changes with increasing reaction temperature. Unlike the CO to CO2 ratios shown in Figure 7a, this ratio magnifies the differences between the effects of various materials of construction on the product selectivity. As the reaction temperature increases, the selectivity to the COx’s increases at a higher rate than the selectivity to the partial oxidation products, although it approaches a constant value for all of the materials except when the empty reactor is used. This suggests that a quasi steady state exists between the rate of generation of the partial oxidation products and the nonselective COx’s that does not occur with the empty reactor where homogeneous reactions are dominant. 4.2. Effect of Metal Frit and Glass Wool Materials. Test quantities of catalysts are typically held in place in laboratory-scale fixed-bed reactor tubes using


Figure 9. Comparison of oxygen and 1,3-butadiene conversion versus reaction temperature obtained for an empty reactor to that for an empty reactor containing glass wool. Feed composition: 8.7% C4H6, 9.7% O2, 9.1% N2, balance of He. Flow rate ) 15 mL/ min (NTP). MARS-3 reactor 6.

Figure 8. Experimental and model-predicted values for the reactant conversion versus temperature using an empty reactor without a frit (filled symbols) and an empty reactor with a frit (empty symbols): (a) oxygen conversion versus reaction temperature; (b) butadiene conversion versus reaction temperature. MARS-2 reactor 6. Feed composition: 9% C4H6, 10% O2, 9% N2, balance of He. Flow rates ) 30 and 90 mL/min (NTP).

various methods, such as porous metal or porous glass frits, glass wool, and other types of materials. These are typically assumed to exhibit an inert behavior in the reaction environment. A brief study was performed to test this hypothesis using both of the MARS reactor configurations. The MARS-2 U-tube reactors employ a type 316 SS porous frit as the catalyst bed support material where the average mesh opening was 20 µm and the frit thickness was 0.76 mm. However, type 316 SS chips were shown above to exhibit significant catalytic activity for 1,3-butadiene oxidation. Consequently, it might be expected that the porous frit could exhibit catalytic activity under typical reaction conditions. On the contrary, the MARS-3 straight-through-tube design uses glass wool to maintain the catalyst in position (see Figure 8 in part 1 of this series of papers). A series of experiments were performed in which a butadiene-rich feed gas was introduced into an empty MARS-2 U-tube reactor either with or without a frit to assess the impact of the frit on the catalytic activity. A similar series of experiments were performed in which glass wool was packed into the empty MARS-3 straight-through-tube design. 4.2.1. Effect of the Metal Frit in the MARS-2 Design. Figure 8 compares the oxygen conversion versus temperature data (Figure 8a) with the 1,3butadiene conversion versus temperature data (Figure 8b) for an empty MARS-2 U-tube reactor without a frit (filled symbols) to an empty MARS-2 U-tube reactor with a frit (filled symbols). In this case, tube no. 6 in MARS-2 was used. The solid lines are the predictions

from a simplified kinetic model, which is described in the next section. The results show that the frit acts as an additional catalytic zone because the conversions of both reactants are higher when the frit is present when compared to the situation where the frit is absent. At the highest temperature used (470 °C), the oxygen conversion approaches 80% at the lowest flow rate (30 mL/min) or the longest contact time for the empty reactor. 4.2.2. Effect of the Glass Wool in the MARS-3 Design. A test of the catalytic activity of glass wool as a catalyst bed support material for the MARS-3 straightthrough-tube reactor is shown in Figure 9. Here, the oxygen and 1,3-butadiene conversions obtained using an empty reactor are compared to those obtained using an empty reactor containing glass wool. The results show that the presence of glass wool has no measurable effect on the observed conversion versus temperature data for either reactant. The use of this material as an inert catalyst bed support is justified. 4.3. Determination of Apparent Kinetics. A knowledge of the contribution of homogeneous gas-phase and wall-catalyzed kinetics to the overall observed kinetics in the presence of a solid heterogeneous catalyst is required before the catalyst performance data can be properly interpreted. A simplified kinetic analysis is given below using the performance data from the empty reactor, reactors that are filled with various materials of construction, and reactors containing metal frits or glass wool catalyst bed support material. 4.3.1. Reactor Metal and Glass Materials. The 1,3butadiene and oxygen conversion versus temperature data shown above in parts a and b of Figure 5, respectively, for both the MARS-3 empty reactor and the one filled with various materials of construction, can be used to obtain basic information on the reaction kinetics for the disappearance of these two reactants. If both 1,3-butadiene and oxygen react according to homogeneous gas-phase and heterogeneous gas-solid catalyzed pathways, then the reaction stoichiometry and reaction rates can be expressed in terms of the following expressions:

a1A + b1B f P1

rA,g ) k1PAR1PBβ1 (homogeneous gas phase) (4)

a2A + b2B f P2

rA,s ) k2PAR2PBβ2 (heterogeneous gas-solid) (5)

In these equations, A and B denote oxygen and 1,3butadiene, respectively, P denotes the reaction products, and the subscripts g and s denote the reaction types,

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i.e., gas-phase and gas-solid catalyzed. It is assumed that the reaction order for A and B for a given reaction can be described by Ri and βi, where i ) 1 or 2. For the homogeneous gas-phase combustion of n-butane, it was determined that R1 ) 1.6 and β1 ) 0.15.23 Determination of the reaction rate constants and reaction orders in eqs 4 and 5 would require independent kinetic data for the homogeneous gas-phase and the heterogeneous gas-solid reactions. However, both of these reactions are occurring in parallel so that parameter estimation is complicated. The energies of activation for these two types of reactions would most likely have significantly different values that would exhibit different sensitivity over a large temperature range (e.g., 300-500 °C). By first assuming that a single reaction occurs over the entire temperature range, it might be possible to identify the temperature range where the particular reactions are dominant. The simplest model that can be used to describe the reactor performance is based upon the following assumptions: (1) butadiene and oxygen both react according to pseudo-first-order kinetics, (2) the gas-phase flow pattern follows plug-flow behavior without a significant molar expansion effect, and (3) the reaction volume zone is constant over the entire range of reaction temperatures. With these assumptions, the integral reactor performance equation becomes24

ln

(

)

( )

-Ea 1 ) kτ ) k0 exp τ 1-X RT

(6)

In eq 6, X is the conversion of either oxygen or 1,3butadiene, k is the first-order rate constant (min-1), τ is the mean residence time based upon the reactor void volume [τ ) V/Q (min)], k0 is the frequency factor (min-1), Ea is the apparent energy of activation (J‚mol-1), T is the reaction temperature (K), and R is the universal gas law constant (8.231 J‚mol-1‚K-1). This model lumps the two parallel reaction pathways defined by eqs 4 and 5 into a single reaction, so it does not distinguish between homogeneous gas-phase and heterogeneous gas-solid reactions. Figure 10 compares the experimental and modelpredicted values for the conversion of either oxygen (Figure 10a) or 1,3-butadiene (Figure 10b), where the latter were obtained by regression using the linearized form of eq 6. These data were obtained using reactor 4 in MARS-3 at contact times (τ ) V/Q) of 1, 2, and 4 s. Data were simultaneously collected for reactors 5 and 6 and exhibited the same type of behavior. The oxygen conversion data provide the most critical test of the single-pathway lumped model because it spans 2-100%. Because the butadiene conversion data only achieve a maximum of ca. 20% at a contact time of 4 s, it is not surprising that they are adequately described by the single-pathway lumped model over this limited range of conversion. However, the lack of agreement between the experimental and model-predicted values for the oxygen conversion over the entire temperature range suggests that a more complicated model is required. This is discussed in more detail in a later paragraph. Inspection of the oxygen conversion model predictions in Figure 10a shows that it overpredicts the experimental data at an intermediate temperature range but underpredicts the data at the higher temperature range. This functional behavior suggests that use of a dualreaction pathway model would provide better agreement

Figure 10. Experimental and model-predicted values for the reactant conversion versus temperature: (a) oxygen conversion versus reaction temperature; (b) butadiene conversion versus reaction temperature. Empty MARS-3 reactor 4. Feed composition: 9% C4H6, 10% O2, 9% N2, balance of He. Flow rates ) 15, 30, and 60 mL/min (NTP). Table 4. k0 and Ea/R Values Extracted from the Oxygen Conversion Data for the Empty Reactor and Reactors Filled with Various Materials of Construction reactor

content

Ea/R, K-1

k0, min-1

4 5 6 4 5 6 4 5 6 4 5 6

empty empty empty type 316 SS type 316 SS type 316 SS titanium titanium titanium glass beads glass beads glass beads

12 718 11 812 11 853 25 874 20 858 23 724 10 330 10 511 10 443 11 791 12 386 11 491

19.4 18.3 18.2 41.3 33.7 38.0 15.6 16.2 15.9 17.5 18.6 17.5

with the experimental data because the heterogeneous gas-solid reaction would be dominant over the low-tointermediate temperature range, while the homogeneous gas-phase reaction would be dominant over the intermediate-to-high temperature range. Two energies of activation would be required to describe these two parallel reaction pathways, as suggested by the reactions described by eqs 4 and 5. Table 4 compares the k0 and Ea/R values that were extracted from the oxygen conversion data from both the empty reactor and the reactors filled with various materials of construction. These are based upon data from reactors 4-6 of MARS-3. Inspection of the k0 and Ea/R values shows that the results for the glass beads and the empty reactors are essentially identical within the experimental measurement error. The corresponding values for titanium are slightly less than those for the reactor filled with glass beads and the empty reactor, while those for type 316 SS are ca. twice the titanium values. The magnitudes of all of the Ea/R values suggest that the intrinsic kinetic resistance is controlling when compared to the external gas-solid

Ind. Eng. Chem. Res., Vol. 44, No. 16, 2005 6463 Table 5. Activation Energies and Frequency Factors Derived from Oxygen Conversion Data in Empty MARS-2 Reactors with and without a Frit 30 mL/min (NTP) reactor

config.

4 4 5 5 6 6

with frit without frit with frit without frit with frit without frit

90 mL/min (NTP)

Ea/R, K-1 k0, min-1 Ea/R, K-1 k0, min-1 14 756 14 732 17 788 18 968 15 738 15 636

19.1 18.7 25.3 26.7 21.4 20.8

14 701 15 485 15 567 14 797 14 913 14 858

17.9 18.8 20.6 19.4 19.2 18.8

Table 6. Activation Energies and Frequency Factors Derived from Butadiene Conversion Data in Empty MARS-2 Reactors with and without a Frit 30 mL/min (NTP) reactor

config.

4 4 5 5 6 6

with frit without frit with frit without frit with frit without frit

90 mL/min (NTP)

Ea/R, K-1 k0, min-1 Ea/R, K-1 k0, min-1 13474 12062 12086 12077 13635 12120

16.2 13.9 15.2 14.9 17.1 14.6

11520 12736 12434 13378 14981 14536

12.6 14.1 14.8 15.9 18.1 17.1

mass transfer. The Ea/R values for type 316 SS are comparable to those for typical selective catalytic pathways so that their effect on the observed rate cannot be neglected. 4.3.2. Reactor Metal Frit Material. The oxygen conversion and 1,3-butadiene conversion versus temperature data shown in Figure 8a,b for the two modes of empty MARS-2 reactor tube operation (i.e., with a frit and without a frit) can be used to extract the apparent Damko¨hler number Da1 ) kτ ) k(Va/Qg) where τ ) Va/Qg is the apparent contact time. The working equation is given as follows where pseudo-first-order behavior and a plug-flow gas have been assumed for the empty tube reactor.

ln

(

)

Figure 11. Axial temperature profiles along the MARS-3 empty reactor centerline (reactor 3). Conditions: 380 °C (top line), 340 °C (middle line), 300 °C (bottom line); flow rate ) 30 mL/min (NTP); feed composition of 9% C4H6, 10% O2, 9% N2, and 72% He (balance).

1 ) k(τe + τf) ) 1-X -Ea -Ea (τe + τf) = k0 exp τ (7) k0 exp RT RT e

( )

( )

In eq 7, τe is the mean residence time in the empty tube (τe ) Ve/Qg) and τf stands for the mean residence time in the porous frit (τf ) Vf/Qg). Because τe . τf, as a result of the significant voidage in empty tube versus the porous frit voidage, eq 7 can be used to extract the energy of activation and frequency factor for the two modes of empty tube operation. Figure 8 compares the experimental and modelpredicted conversion versus temperature results when the data are fitted to eq 7. A summary of the frequency factors and energies of activation that were derived using these data is given in Tables 5 and 6. The presence of the frit induces additional catalytic activity that becomes significant with increasing temperature, particularly when the temperature exceeds ca. 420 °C. This suggests that other catalyst bed support materials should be utilized in any future work. 4.4. Temperature Profiles in Empty Reactor Tubes. A precise determination of the contact time for empty tube operation in MARS-2 at a given temperature is not possible. This is because the length of the active catalytic zone is dependent upon the reaction temperature owing to the immersion of the reactor U-tube into the fluidized-bed sand bath heater. As the sand bath

temperature is increased, the effective length of the reaction zone also increases because of the higher heattransfer rate from the sand bath to the entire U-tube. However, the method used for heating the MARS-3 reactor is different from the one used for MARS-2 because it is based upon enclosing the reactor tube inside a heated aluminum block (see part 1 of this series of papers). As a result of this design, the lengths of the heated zones at the reactor inlet and outlet are shorter when compared to the MARS-2 design. The presence of the frit in the MARS-2 also increases the apparent contact time, but it is not possible to determine the extent of the increase with the available data. This would require a carefully executed pulse or step input tracer response experiment, but it was not performed as part of this study. 4.4.1. Effect of the Set Point at a Constant Inlet Gas Flow Rate. Insight into the lengths of these heated zones at a given reactor temperature set point can be obtained by inspection of Figure 11, which shows the axial temperature profile along the MARS-3 reactor centerline. These data were measured by sliding a 1.5875-mm (1/16 in.) o.d. type J thermocouple along the reactor centerline, where it was maintained in position using a disk with several spokes. The material between the spokes was removed so that the gas flow was not restricted. The results show that the temperature of the reactor increases from the cooler inlet temperature to the desired set point following nearly the same temperature gradient (i.e., dT/dz is constant) for each set point. About 20% of the overall reactor tube length is required to approach the desired set point, compared to less than 5% of the tube length at the reactor outlet. It demonstrates that extraction of the true contact time for the reaction mixture clearly requires detailed data on the reactor temperature profiles. If the entire reactor tube was assumed to exist at the desired set-point temperature, the contact time would be larger than the actual one. Table 7 compares the reactor set-point temperature at two different gas flow rates to the average reactor temperature where the latter was obtained by integration of the temperature profile data shown above in Figure 11. It shows that the mean temperature is always less than the actual set point and the temperature difference increases with increasing gas flow rates. Collectively, the temperature profile data shown in Figure 11 and the average values given in Table 7 suggest that any kinetic model based upon empty reactor tube performance data requires both microscopic mass and energy balances. Although the temperature

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Table 7. Comparison of Reactor Set-Point Temperatures with the Average Temperatures for MARS-3a flow rate, T average feed composition, % mL/min set point, T, ∆T, (NTP) °C °C °C C4H6 O2 He H2O 30 60 30 60 30 60 30

300 300 340 340 380 380 380

285.2 278.5 319.0 317.1 356.5 356.3 356.1

-14.8 -21.5 -21.0 -22.9 -23.5 -23.7 -23.9

8.6 8.6 8.6 8.6 8.6 8.6 8.0

10 10 10 10 10 10 9.3

89.7 89.7 89.7 89.7 89.7 89.7 74.3

0 0 0 0 0 0 8.4

a The latter were obtained by integration of the temperature profiles.

point is incorrect, and this effect must be accounted for in any detailed reactor kinetic modeling. 4.4.3. Effect of Steam Addition. Figure 13 is based upon the same temperature difference concept explained above, except it shows the effect of steam addition on the axial temperature profiles at a fixed flow rate of 30 mL/min (NTP). Here, the difference is defined as ∆T ) Tsteam - Tno steam at a reactor set point of 380 °C. The addition of steam causes the temperature difference to be positive at the reactor inlet, but the difference is nearly constant at a slightly negative value over most of the inner section of the reactor tube. The higher heat capacity of the gas mixture in the presence of steam could be one explanation for the latter observation. 5. Summary and Conclusions

Figure 12. Difference between the axial temperature profiles along the MARS-3 empty reactor centerline (reactor 3) obtained at two different gas flow rates. The difference is defined as ∆T ) T(Q ) 60 mL/min) - T(Q ) 30 mL/min). Conditions: 380 °C (top line), 340 °C (middle line), 300 °C (bottom line); flow rates ) 30 and 60 mL/min (NTP); feed composition of 9% C4H6, 10% O2, 9% N2, and 72% He (balance).

Figure 13. Effect of steam addition on the axial temperature profiles along the MARS-3 empty reactor centerline (reactor 3). The difference is defined as ∆T ) Tsteam - Tno steam. Conditions: T set point ) 380 °C; flow rate ) 30 mL/min (NTP); feed composition of 9% C4H6, 10% O2, 9% N2, 20% H2O, and 52% He (balance).

profile data were not collected for the case where the central zone of the reactor tube was filled with catalyst particles, the effective thermal conductivity of the catalyst zone would not have a significant impact on the temperature profile on either side of this zone. Hence, the temperature profile measured using an empty tube would also be a good approximation to the situation where catalyst was present in the tube. 4.4.2. Effect of the Gas Flow Rate. The effect of increasing the gas flow rate on the temperature profiles is shown in Figure 12 as a temperature difference defined as ∆T ) T(Q ) 60 mL/min) - T(Q ) 30 mL/ min). The three lines correspond to different reactor set points, namely, 300, 340, and 380 °C. The effect of the flow rate on the observed temperature difference becomes less important as the reactor set point is increased. To assume that the gas flow rate has no effect on the average temperature at a given temperature set

The gas-phase oxidation of 1,3-butadiene was studied in two different parallel laboratory-scale reactor designs to obtain an independent assessment of the contribution of homogeneous gas-phase and wall-catalyzed reactions that were suspected to occur during regular catalyst performance evaluation tests. The reactor designs included classical U-tube and conventional straightthrough-tube configurations that were heated by individual fluidized-bed sand baths and split-tube furnaces, respectively. These reactor configurations were part of a parallel reactor system called MARS. MARS was designed for detailed performance evaluations of solidcatalyzed gas-phase partial oxidation chemistries using a parallel arrangement of six fixed-bed microreactors and automation for safe, unattended operation. Although the parallel evaluation of heterogeneous catalyst libraries is now a widely practiced technology, the issue of reactor system inertness has not been studied in any detail for specific applications. Gas-phase partial oxidation reactions provide special challenges in the creation of an inert reactor system environment that is external to the catalyst zone because empty volume heated zones can lead to undesired homogeneous reactions. In addition, commonly used materials of construction for fabrication of reactor system components can promote other undesired reactions caused by contact of the reaction mixture with heated metal and other types of surfaces. The reaction products generated by these homogeneous gas-phase and wall-catalyzed reactions can also be independently generated in parallel from the solid catalyst under investigation. These two independent sources of the same undesired reaction products can lead to incorrect conclusions on the catalyst performance. Generally, the presence of homogeneous and wall-catalyzed reactions results in the additional generation of partial and total combustion products that lower the overall yield of the desired product. Experiments in which reaction mixtures containing 1,3-butadiene and oxygen were contacted in the U-tube and straight-through-tube configurations showed that the U-tube microreactor design produced a higher conversion of the reactants at the same reaction conditions. The reaction products were mainly COx’s, although minor amounts of organic reaction products, such as acetone and furan, were also detected. The higher conversions associated with the U-tube reactor when compared to the straight-through-tube design can be attributed to a longer heated section of the U-tube that also results in a greater net contact time. The straight-through-tube reactor design with the split-tube furnace for heating allows the length of the pre- and


postheated zones to be controlled at a lower temperature where nonselective reactions exhibit lower activity. The ability to precisely control the temperature profiles in the catalytic as well as the pre- and postheated zones is an important feature for reducing the contribution of undesired gas-phase and wall-catalyzed reactions. The addition of various materials of construction to each reactor tube was also performed and allowed an independent assessment of whether a particular material induced significant wall-catalyzed reactions. For both reactor configurations, the order of reactivity as measured by the increased oxygen conversion was glass beads < glass wool < Ti < type 316 SS. While glass is the most inert material for this application, it also has the least favorable heat-transfer characteristics and lowest mechanical strength to contain a possible detonation. These results suggest that special-purpose glasslined tubing and other inert coatings would reduce the contribution of wall-catalyzed reactions. In addition, careful design and temperature control of heated open volumes must also be considered when testing actual heterogeneous catalysts to minimize the contribution of homogeneous gas-phase reactions. Acknowledgment The authors thank Leo E. Manzer in DuPont CR&D for his leadership and insight that was provided on the catalysis and process comparisons for both the 1,3butadiene and n-butane oxidation routes to THF. Literature Cited (1) Stadig, W. Three Inventions Combine to Yield New Route to THF. New Technology May Also Find Application in Other Processes. Chem. Process. 1992, 8 (Aug), 27. (2) Mills, P. L.; Manzer, L. E.; Roy, S. D. Green Chemistry Process Alternatives for THF Manufacture. AIChE 2003 Annual Meeting, Session 551 on Green Chemistry and Reaction Engineering III, Nov 18, 2003; Paper 551f. (3) Contractor, R. M. Dupont’s CFB Technology for Maleic Anhydride. Chem. Eng. Sci. 1999, 54 (22), 5627. (4) Contractor, R. M.; Letts, W. J. (E. I. DuPont de Nemours and Co., Wilmington, DE). Process for Manufacture and Use of Improved Attrition Resistant Catalyst. U.S. Patent 6,107,238, Aug 22, 2000. (5) Centi, G.; Cavani, F.; Trifiro, F. Selective Oxidation by Heterogeneous Catalysis. Fundamental and Applied Catalysis; Kluwer Academic/Plenum: New York, 2001. (6) Barteau, M. A.; Lopez, L.; Buttrey, D. J. In Situ Studies of the Mars-van-Krevelen Mechanism in Hydrocarbon Oxidation. AIChE 2004 Annual Meeting, Austin, TX, Nov 2004. (7) Bither, T. A., Jr.; McClellan, W. R. (E. I. DuPont de Nemours and Co., Wilmington, DE). Preparation of Furan. U.S. Patent 4,322,358, Mar 30, 1982. (8) Contractor, R. M.; Horowitz, H. S.; Sisler, G. M. (E. I. DuPont de Nemours and Co., Wilmington, DE). Vapor Phase Catalytic Oxidation of n-Butane to Maleic Anhydride Incorporating in situ Catalyst Calcination/Activation. U.S. Patent 5,519,149, May 21, 1996.

(9) Mahajan, S.; Menzies, W. R.; Albright, L. F. Partial Oxidation of Light Hydrocarbons. 1. Major Differences Noted in Various Tubular Reactors. Ind. Eng. Chem. Process Des. Dev. 1977, 16 (3), 271. (10) Mahajan, S.; Menzies, W. R.; Albright, L. F. Partial Oxidation of Light Hydrocarbons. 2. New Techniques for Investigating Surface Reactions. Ind. Eng. Chem. Process Des. Dev. 1977, 16 (3), 275. (11) Mahajan, S.; Albright, L. F. Partial Oxidation of Light Hydrocarbons. 3. Mechanism Incorporating Key Surface Reactions with Gas-phase Steps. Ind. Eng. Chem. Process Des. Dev. 1977, 16 (3), 279. (12) Schlatter, J. C., Ed. A Practical Guide to Catalyst Testing; Catalytic Study No. 4186 TE; Catalytica Press: Mountain View, CA, 1987. (13) Rodemerck, U.; Baerns, M. Tools for High-Throughput Experimentation in the Development of Heterogeneous Catalysts; Springer Series in Chemical Physics; Springer-Verlag: Berlin, 2004; p 259. (14) Schuth, F.; Busch, O.; Hoffman, C.; Johann, T.; Kiener, C.; Demuth, D.; Klein, J.; Schunk, S.; Strehlau, W.; Zech, T. Highthroughput Experimentation in Oxidation Catalysis. Top. Catal. 2002, 21 (1-3), 55. (15) Kourtakis, K. (E. I. DuPont de Nemours and Co., Wilmington, DE). Cation and Vanadium Substituted Heteropolyacid Catalysts for Vapor Phase Oxidation. U.S. Patent 5,510,308, Apr 23, 1996. (16) Kourtakis, K.; Sonnichsen, G. C. (E. I. DuPont de Nemours and Co., Wilmington, DE). Catalyst and Method for Vapor Phase Oxidation of Alkane Hydrocarbons. U.S. Patent 5,543,532, Aug 6, 1996. (17) Kourtakis, K.; Sullivan, J. D. (E. I. DuPont de Nemours and Co., Wilmington, DE). Molybdenum Based Oxidation Catalysts. U.S. Patent 6,500,779, Dec 31, 2002. (18) Maier, W. F. Combinatorial Chemistry in Catalysis and Materials Research. Chem. Ing. Tech. 2003, 75 (8), 1048. (19) Laskin, A.; Wang, H.; Law, C. K. Detailed Kinetic Modeling of 1,3-Butadiene Oxidation at High Temperatures. Int. J. Chem. Kinet. 2000, 32 (10), 589. (20) Blackstone, C. M.; Lerou, J. J. An Automated Facility for Catalyst and Process Research. Het Ingenieursblad 1986, 55 (6), 289-294. (21) Mills, P. L.; Nicole, J. F. Multidimensional GC Method for Analysis of C4 Hydrocarbon Oxidation Reaction Products. 1. Detailed Comparison of Single and Multiple Oven Methods. Ind. Eng. Chem. Res. 2005, in preparation. (22) Moulijn, J. A.; Perez-Ramirez, J.; Berger, R. J.; Hamminga, G.; Mul, G.; Kapteijn, F. High-Throughout Experimentation in Catalyst Testing and in Kinetic Studies for Heterogeneous Catalysis. Catal. Today 2003, 81, 457-471. (23) Flagan, R. F.; Seinfeld, J. H. Fundamentals of Air Pollution; Prentice Hall: Englewood Cliffs, NJ, 1998. (24) Froment, G. F.; Bischoff, K. B. Fixed bed reactors. Chemical Reactor Analysis and Design; Wiley-Interscience: New York, 1990; pp 125-197.

Received for review December 2, 2004 Revised manuscript received May 31, 2005 Accepted June 1, 2005 IE048830Z

Multiple Automated Reactor Systems (MARS). 2. Effect of Microreactor

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