Systematic Regeneration of Waste Sulfuric Acid in Semiconductor

Article pubs.acs.org/IECR

Systematic Regeneration of Waste Sulfuric Acid in Semiconductor Manufacturing Using Batch Vacuum Distillation Seongho Park,† Jaeheum Jung,† Kiwook Song,† Krishnadash S. Kshetrimayum,† Changhyun Jeong,‡ and Chonghun Han*,† †

School of Chemical and Biological Engineering, Seoul National University, Gwanak-ro 599, Gwanak-gu, Seoul 151-742, South Korea Basic Engineering Department, Samsung Engineering Co. Ltd., Samsung GEC, 26, Sangil-ro 6-gil, Gangdong-gu, Seoul,134-728, South Korea

‡

S Supporting Information *

ABSTRACT: We describe herein a systematic regeneration of waste sulfuric acid produced in semiconductor manufacturing, using batch vacuum distillation (BVD). During the recycling process, dilute sulfuric acid feed was continuously concentrated and fed back to the original wafer washing step. It consisted of a batch tank to charge the feed solution, condenser to capture generated vapor, receiving tank to receive condensed distillate liquid, and vacuum pump to reduce the system pressure. The improper control of the vacuum operation led to incomplete condensation; consequently, the vacuum pump became dysfunctional. The goal of this study was to prevent such mishap. After the feed condition was defined, a basic design was conceived, and the main characteristics of the BVD were determined. The results of sensitivity analyses on the feed and operating conditions have been discussed. The strategies for designing the vacuum pump’s capacity should be changed depending on phase equilibria at the target pressure.

1. INTRODUCTION Among the heavy industrial chemicals, sulfuric acid is considered as the most fundamentally important chemical because it has a number of large-scale applications not only within the chemical industry but also in other industries.1 It is one of the most widely consumed chemicals in the world and used as an index of the scale of national chemical industries. Indeed, sulfuric acid is one of the three most extensively used chemicals in inorganic chemical industries.2 In the past, sulfuric acid was most heavily used in manufacturing fertilizers. Other important applications include its use in wastewater processing, metallurgy industries, specialty chemical manufacturing, and petroleum refining. Recently, the semiconductor industry has been a large consumer of sulfuric acid products. Among other processes, it is usually employed in wafer cleaning. The use of sulfuric acid to strip a photoresist from silicon wafers is a widely employed technique in semiconductor manufacturing.3 In most cases, the acid is combined with hydrogen peroxide to oxidize the stripped photoresist material.3 Moreover, the regeneration of waste acid has been studied in many fields to improve the sustainability of the industrial processes. Therefore, the recovery of waste sulfuric acid in the semiconductor industry has also become an important issue from environmental and economic standpoints. Several examples of recycling waste sulfuric acid have been reported previously.4−11 Although there does not exist a single procedure which could be universally applied for regenerating all types of waste acids, the distillation technologies are usually used for concentrating dilute sulfuric acid. For example, von Plessen et al. developed a process in which dilute sulfuric acid is recycled by distillation in vacuum using a Pauling apparatus.4 Mazzafro et al. described a process for the purification and concentration of sulfuric acid containing nitric acid byproducts © 2014 American Chemical Society

and organic components. They used a stripping column wherein the contaminants were removed followed by a multiple-effect vacuum concentration system, which regenerated sulfuric acid by evaporating water in a series of steps.9 Chou et al. suggested the recovery and regeneration of sulfuric acid by the simultaneous removal of both water and organic compounds from the spent sulfuric acid catalyst in the alkylation of olefins and alkanes.10 Moreover, there are academic research papers on the regeneration of sulfuric acid. Tongwen et al. conducted a series of experiments for the recovery of sulfuric acid from titanium white waste liquor.12 In their study, the effects of some important factors such as ion exchange capacity, the content of benzylhalogen, and relative compositions of the liquor were experimentally investigated using the membrane dialysis process.12 Jeong et al. applied this diffusion method involving anion exchange to recover sulfuric acid from the waste solution produced in diamond manufacturing.13 Although several industrial applications and academic findings on either the batch systems or vacuum pumps are known, not much attention has been paid toward combining the two systems. Distefano14 developed a rigorous mathematical model involving a multicomponent batch distillation process. E. R. Robinson15 calculated the optimal reflux policies for a batch distillation process, and Sigurd et al.16 proposed a model of a multivessel batch distillation (MVBD) column. Recently, a nonlinear model-based control algorithm was proposed for a simulated batch reactive rectifier by Prateek et Received: Revised: Accepted: Published: 8543

January 2, 2014 April 16, 2014 April 25, 2014 April 25, 2014 dx.doi.org/10.1021/ie500016v | Ind. Eng. Chem. Res. 2014, 53, 8543−8552

Industrial & Engineering Chemistry Research

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Figure 1. Schematic of sulfuric acid and hydrogen peroxide mixture treatment.

al.17 Lin et al. studied a startup modeling problem of batch distillation starting from an empty cold state.18 Several types of vacuum pumps and their operational principles have been either claimed in various patents or published in academic papers. Canizares et al. carried out an evaluation of a simple batch distillation process involving a vacuum pump for treating effluents generated by metalworking factories.19 Houska et al.20 developed a mathematical model of a vacuum cooling system and validated it through several experiments on fruit/sugar mixtures and water systems. In particular, the latter was not based on the batch distillation processes but rather on the simple vacuum cooling systems. To the best of our knowledge, however, no academic research on the regeneration of waste sulfuric acid produced in the semiconductor industry with regard to process system engineering has been conducted hitherto. Most of the research dealing with the recycling of sulfuric acid has focused on the problem in either one or more of the following ways: (i) how to make the recycling process compatible with the existing processes; (ii) how to develop recovery methods, which circumvent the existing patents; and (iii) how to construct devices relevant for the regeneration process without considering equilibrium principles, kinetics, thermodynamic properties, and optimization. However, in this study, we aimed to develop a sulfuric acid regeneration utility for the semiconductor manufacturing industry based on process design principles. By considering the characteristics of upstream and downstream processes, optimal designs of recycling systems have been developed.

acid, is drained out after the cleaning process. We developed a recycling system to reuse this discarded liquid, which can be considered as our feed solution. SPM reaction chemistry can be used to evaluate the feed conditions because the product of the reaction becomes the feed material in the regeneration process. According to the literature,21 when H2O2(aq) is brought in contact with H2SO4, two equilibrium reactions occur: H2O2(aq) protonation equilibrium (eq 1) and redox equilibrium transformation of Caro’s acid (eq 2). H 2O2 (aq) + H3O+(aq) ↔ H3O+2 (aq) + H 2O ° = 4.3 ± 1.2 kcal/mol Δr H298

(1)

H 2O2 (aq) + HSO−4 (aq) ↔ HSO−5 (aq) + H 2O ° = −2.6 kcal/mol Δr H298

(2)

Thus, it can be deduced that at high concentrations of H2SO4, both the reactions converge to eq 1 while the H2O2(aq) protonation equilibrium is almost fully shifted to the right.21 Consequently, the reaction to form Caro’s acid from protonated hydrogen peroxide can be written as H3O+2 (aq) + HSO−4 (aq) = H3O+(aq) + HSO−5 (aq) ° = −6.9 ± 1.2 kcal/mol Δr H298

(3)

Notably, the overall reaction describes the decomposition mechanism of H2O2(aq) to H3O+(aq) in sulfuric acid solution and that its heat of reaction is weakly exothermic. In reality, however, an enormous amount of heat is emitted during the SPM chemistry. Therefore, the heat of dilution (or mixing) of H2SO4 in water should also be taken into account.

2. PROCESS DESCRIPTION 2.1. Feed Condition. In the semiconductor industry, a sulfuric acid and hydrogen peroxide mixture (SPM) is usually introduced in the wafer cleaning processes. High reactivity and vigorous heat emission are known to be the main characteristics of the solution, which make it possible to remove either organic particles or metal impurities from the silicon wafers. As shown in Figure 1, sulfuric acid and hydrogen peroxide are mixed and then injected into the wafer, wherein the cleaning process occurs. After undergoing washing for a while, the cleaned wafer goes through other downstream processes, and another contaminated wafer is introduced into the reactor for SPM treatment. The effluent, typically containing diluted sulfuric

H 2O(l) + H 2SO4 (l) = H3O+(aq) + HSO−4 (aq)

(4)

Comparison calculation of the heat emissions of the two reactions were conducted, and the results are represented in Table S1 in the Supporting Information. The specific enthalpy of each compound was calculated by simulator (Aspen Plus V7.3) using the thermodynamic properties of the symmetric electrolyte NRTL (nonrandom two-liquid model). Consequently, the heat of dilution of sulfuric acid (eq 4) in water was about 3 orders of magnitude larger than that in eq 3. Therefore, 8544

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2.2. Objective and Constraints. In this study, we aimed to obtain a feasible regeneration system for the dilute sulfuric acid for each feed condition defined above. In order to make the design criterion definite, in all cases, the volume of the feed solution and its treatment time were taken as 100 L and 1 h, respectively. The target purity was set as 93 wt % H2SO4. Among many separation techniques, distillation was chosen in our study because the difference in the boiling points of H2SO4 and H2O is large. Moreover, batch distillation seemed more suitable for our system compared to continuous distillation because it performed significantly better in terms of flexibility, operability, and versatility. In general, it is wellknown that the batch distillation has the advantages of versatility and flexibility in treating different mixtures.22 The process design should have inherent flexibility; for instance, in our case, the feed condition changed periodically. There is a strict constraint on materials when dealing with sulfuric acid because only a few materials are resistant to the corrosiveness of the hot acid. Increased quantity of the toxic, sulfur trioxide gas is generated as the system temperature increases. In addition, from the viewpoint of recycling, it is better for the resulting concentrated sulfuric acid to be not too hot because its initial temperature is already assumed to be 25 °C. Therefore, it is more desirable to return the sulfuric acid to its original state because the downstream process (in this case SPM reaction) is already characterized by the low-temperature feed solution. For these reasons, we kept the distillation temperature as low as possible. Thus, an appropriate vacuum condition was chosen as the operating pressure because in essence, the boiling point of the solution decreased as the pressure decreased. We also found that it was a common practice to distill the sulfuric acid at lower pressures. Hence, we decided to reduce the pressure with a vacuum pump, which is usually used to create a low-pressure system. The limitation of the cooling utility should also be taken into consideration when it comes to the semiconductor manufacturing processes. Owing to the nature of this industry, which is different from that of other bulk chemical manufacturing industries, it is likely that the cooling capacity of the plant is not adequate to condense all the vapor generated during the distillation process. Additionally, it may be difficult to design an enormous condenser, which is needed to capture the hot vapor. From the viewpoint of semiconductor manufacturing, it would be challenging to make a large condenser compatible with the original processes because such condensers are usually small and part of several types of compact batch distillation equipment. Despite the aforementioned two constraints (cooling utilities and layout limit), the total condensation must be guaranteed for the condenser because it is crucial that the vacuum pump be kept free from corrosion and heavy suction load. Sulfuric acid mist particles in the vapor are certainly critical to the vacuum pump; therefore, appropriate pressure control and vapor treatment are necessary to capture the particles through the condenser. Therefore, when designing a condenser for acid regeneration using vacuum distillation, particular attention should be heeded to the objective and constraints described hereinabove. In summary, we selected the BVD for sulfuric acid regeneration. Owing to the nature of the original process, flexibility and low temperature were required. We employed a still pot charged with the feed liquid, condenser to condense the generated vapor, receiving tank to collect the condensed distillate, and vacuum pump to lower the pressure by pumping

we concluded that a major portion of the heat evolved was because of the mixing of sulfuric acid and water. In other words, when sulfuric acid and hydrogen peroxide were mixed together in an aqueous solution, the latter was decomposed into water in the form of hydronium ions, which were further absorbed by H2SO4 liberating a huge amount of heat. The reactive oxygen species vigorously dissolved and oxidized any organic or inorganic impurities in the Caro’s acid environment. In entirety, the so-called Piranha solution chemistry can be written as follows: H 2SO4 + H 2O2 → H3O+ + HSO−4 + O(radical)

(5)

The Piranha solution is usually prepared by mixing sulfuric acid and hydrogen peroxide in a particular volumetric ratio. A commonly used mixing ratio of H2SO4 and H2O2 falls over the range of 2:1 to 7:1. While H2O2 acts as an actual oxidizing agent, it is usually included as a limiting reactant. Figure 2 depicts the estimation of temperatures and compositions at different volumetric mixing ratios. The

Figure 2. Estimated feed conditions: temperature and composition.

reactants were assumed as 96 wt % H2SO4 and 30 wt % H2O2 in an aqueous solution at 25 °C, 1 atm. In mass balance, all H2O2 species were assumed to be decomposed to H2O and oxygen radicals. In particular, it was also assumed that the latter was removed from the system in the form of other compounds such as CO2. Thus, only H2O and H2SO4 species were implemented in the energy balance calculation. In the RStoic reactor model of the simulator, H2O and H2SO4 were incorporated with their respective ionization reactions 4. The conversion was assumed to be 100% because sulfuric acid was one of the strongest electrolytes and its concentration was high in this system. Reaction 3 was not considered here because, as mentioned earlier, the heat of reaction was smaller compared to that of that in reaction 4. Consequently, as the volumetric mixing ratio increased, the product’s temperature decreased, and the concentration of sulfuric acid remained high. This is because at high mixing ratios, the relative amount of limiting reactant H2O2 became small, which made the reaction system less exothermic. Hereinafter, the feed is defined as the diluted product of the Piranha reaction. For each mixing ratio, the product condition is listed in Table S2 in the Supporting Information. It only consists of the binary compounds of H2SO4 and H2O, and any other chemicals are assumed to have been removed before entering the recycling system. 8545



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Unfortunately, however, the target operating pressure could not be reduced indefinitely because of the limited cooling by the condenser. As an example, Figure 5 represents the vapor−

noncondensable gas. The schematic of the overall process is shown in Figure 3. Next, the specific operating conditions of BVD were determined to obtain 93 wt % H2SO4.

Figure 5. Equilibrium curve of pure water calculated using electrolyte NRTL (nonrandom two-liquid) model.

liquid equilibria of pure water. If we assumed that the lower bound of the cooling utility’s temperature was 20 °C and minimum temperature difference in the heat exchanger was 10 °C, the minimum condensation temperature was computed as 30 °C. Thus, we concluded through the phase diagram that the pressure could not be reduced below 42.5 mbar. After determining the target vacuum pressure, we could also determine the operating temperature. For example, if we defined the target pressure as 42.5 mbar, the equilibrium temperature was determined from the phase equilibria of water and sulfuric acid at that pressure. Figure 4 reveals that at 42.5 mbar, 93 wt % H2SO4 liquid is in equilibrium with the vapor containing 15 wt % H2SO4 when the temperature reaches 182.5 °C. In other words, we could establish the operating temperature as the bubble point of 93 wt % H2SO4, at which the solution started to boil. 2.4. Design Basis. It is necessary to specify all the parameters in the process as a basis. In our basic design, we defined the feed by mixing sulfuric acid and hydrogen peroxide in a volumetric ratio of 3:1. From the heat and mass calculations described in section 2.1, the product of the reaction at 164 °C was 80 wt % sulfuric acid solution, which was concentrated to achieve 93 wt %. As mentioned in section 2.2, the volume of the feed is 100 L with the operating duration specified to be 60 min. A rough energy balance indicated that 25 kW of the heat had to be provided to meet the given throughput. Considering the condenser cooling as described in the previous section, we determined the target pressure to be 42.5 mbar. We assumed that the free volume of the system was 200 L, which was two times greater than the volume of the feed. We also assumed the vacuum pump’s capacity (S) to be 50 L/min. Then, the vacuum time was computed to be 12.6 min using the design equation of the vacuum pump, which will be explained later in this article. It took 12.6 min for the pressure to drop from 1013.25 mbar to 42.5 mbar. Notably, this duration was considerably lesser than the entire duration of the process (60 min).

Figure 3. Schematic of batch vacuum distillation.

2.3. Design Selection. Figure 4 is a Txy diagram of the binary system of H2SO4 and H2O at 1013, 300, and 42.5 mbar.

Figure 4. Txy diagram of sulfuric acid and water binary system at different pressures: 1013, 300, and 42.5 mbar.

Evidently, the bubble point curve shifted downward as the pressure decreased; it also decreased at the target purity of 93 wt % with decreasing pressure. This meant that the point at which the concentrated sulfuric acid and water-rich vapor were in equilibrium could be gradually lowered by reducing the pressure. Consequently, the operating temperature could also be lowered, and a BVD process, which satisfied the aforementioned constraints, could be designed. 8546



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p(t ) = p0 e−S / Vt

3. MODELING OF SULFURIC ACID REGENERATION SYSTEM 3.1. Comments on Thermodynamic Properties. It is important to select a relevant thermodynamic model because the water−sulfuric acid system exhibits nonideal behavior. Recently, a model called symmetric electrolyte NRTL23 was developed and applied to the sulfuric acid−water−sulfur trioxide system. This comprehensive model covered the entire concentration range and showed an accurate correlation with the experimental VLE data. Figure 6 shows a comparison between the conventional asymmetric and symmetric electrolyte NRTL models devel-

(7)

where p(t) is the system pressure, which is a function of time t, and p0 is the initial pressure. S is vacuum pump’s capacity whose dimension is expressed as volume per time, and V is free volume of the system. Using the equation above, the pressure of the system could be specified in time coordinates and, hence, the unsteady vacuum operation could be described using the simulation. Additional information such as vapor−liquid equilibria, reboiler/condenser heat duty, system temperature, and vapor emissions could be incorporated into or extracted from the repetitive steady-state simulations. Thus, by applying the simulation methodology described herein, we could derive the dynamic behavior of the system.

4. RESULTS AND DISCUSSION 4.1. Basic Design. A basic design was implemented to understand the characteristics of the vacuum batch distillation process. All parameters used here have been defined and specified in section 2.4. Figure 7 depicts the base case design. The pressure, instantaneous vapor generated in the still pot, condenser duty, temperature, liquid holdup, and liquid composition of sulfuric acid were recorded in the whole time coordinate.

Figure 6. Comparison of Pxy diagram at 150 °C between symmetric and asymmetric electrolyte NRTL (nonrandom two-liquid) models based on previously reported experimental data.6

oped by Chau−Chyun Chen. The experimental data of Pxy at 150 °C was extracted and recalculated using the Perry’s Chemical Engineers’ handbook.24 The symmetric electrolyte NRTL data was generated using a simulator wherein we incorporated all the binary interaction parameters of the regressed electrolyte NRTL and reaction equilibrium constant parameters reported previously.25 Evidently, the symmetric electrolyte NRTL model fitted with the experimental VLE data more accurately compared to the conventional asymmetric model. 3.2. Simulation Methodology. A realistic unsteady dynamics of vacuum distillation in the simple batch tank (not a column) can be described using a steady-state model. The basic premise is that if the process time is divided sequentially, the batch process could be modeled as a succession of discrete steady-state simulation calculations. The repetitive calculations of the continuous distillation model with different feed stream specifications as well as unit conditions could be interpreted as subsequent nonstationary batch operations. The product stream of the previous computation was introduced as an input feed stream for the next simulation. In this way, it was in principle possible to simulate the unsteady-state distillation, if we chose an adequately short interval of time.26 In addition, in this study, a vacuum pump design eq 7 was introduced to describe the characteristics of the vacuum process.

Figure 7. Results of basic design: dynamics of batch vacuum distillation. 8547



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from the vacuum pump operation. The latter, rather than the former, was the main cause of the vapor generation. This rationale could be verified by studying the isobaric period, wherein the pressure remained at target pressure (43.5 mbar) and constant heat (25 kW) was being supplied, and the vapor was generated at a constant rate (5.95 g/s on average). Comparing this rate with the maximum vapor generation rate (16.20 g/s) in the depressurization step, we concluded that in the early stage, the pressure effect rather than the steady heat supply caused a sharp increase in the vapor flow, and therefore a heavier duty was imposed on the heat exchanger. Notably, the intense cooling duty, equivalent to the large heat exchange area, could pose a serious problem if the layout constraint existed as mentioned in section 2.2. Consequently, the cooling duty on the condenser could be reduced by controlling the rate of pressure drop, which in turn could slow down the rate of vapor generation. Now, the focus of the problem was pressure control. By differentiating the vacuum pump eq 7 with respect to time and rearranging it, we obtained the following:

Based on the results of the BVD dynamic system, the process is divided into two parts: a depressurizing step, which is the early stage of the process, and an isobaric period, which is the remaining part of the distillation process. The depressurizing step is defined as the period during which the pressure decreased from the initial value to the final vacuum pressure, whereas the isobaric period is defined as the phase wherein a target low pressure was maintained. In the depressurizing step, the pressure decreased exponentially based on eq 7. As the vacuum pump was in operation, a large volume of vapor was generated suddenly, which imposed a high duty on the condenser. When the subcooled feed reached the bubble point during the depressurization process, the vapor started to evaporate. The system temperature reduced rapidly because of both the latent heat of vaporization effect and the decreased vapor pressure effect. However, in the isobaric period, the remaining liquid whose temperature was reduced in the previous step was heated up gradually. Although the vapor was generated during this stage, its volume was insignificant compared to that of the vapor generated in the depressurizing step. This implies that the cooling capacity needed in the remainder part of the process is not high. Nevertheless, most of the whole process time was devoted toward the isobaric step. This was because the heat loss from the early evaporation was enormous, and a large amount of heat would need to be additionally supplied as compensation. The instantaneous production rate of vapor and its condensing duty are particularly important variables in BVD because they have an enormous effect on the design such as maintenance of the vacuum pump and process layout size. We had mentioned earlier in section 2.2 that the total condensation in the condenser must be guaranteed to protect the vacuum pump from the mist attack. In order to achieve that, the vapor flow, which contained the largest amount of latent heat, should be captured through the heat exchanger. After contact with the condenser, the vapor fraction of that flow should be maintained at zero. Once the largest amount of heat containing vapor is totally condensed, the remainder of the vapor is also captured. Thus, if the heat exchanger is designed to condense the vapor containing the largest amount of heat, it can also guarantee that the remainder of the vapor would be totally condensed. In our case, when the duration was 220 s, the maximum amount of vapor 16.2 g/s and maximum condenser duty −44.63 kW were observed. The simulation results are summarized in Table 1. The vapor generation in the depressurization step was due to both the heat absorption from reboiler and pressure decrease

d ln p S =− dt V

It is evident that the rate of pressure drop is proportional to the capacity of vacuum pump S. The differential eq 8 indicates that the speed of pressure drop could be lowered if a lowperformance vacuum-pump is used. Because we fixed the free volume V as 200 L, only the value of S could be adjusted. Thus, there exists an opportunity to regulate the system pressure by selecting an appropriate capacity of the vacuum pump. 4.2. Sensitivity Analyses of Different Feed Conditions. For each feed condition, the most appropriate processes were designed. To that end, the sensitivity analyses on each feed mixture were conducted to see how the processes responded to the variation in the vacuum pump’s capacity. Figure 8 represents the instantaneous vapor emission of the feeds with the mixing ratios of 2:1 and 7:1. For these two extreme cases, several simulations were carried out. All the variables, except the pump’s capacity, were fixed, and the same as those for the base case. For the mixing ratio of 2:1, the pump’s capacity (S) was changed from 60 L/min to 20 L/min. As seen in Figure 8a, the maximum amount of vapor generated decreased as the value of S decreased. When S was 60 L/min, the maximum vapor production rate was 22.53 g/s. This rate was approximately 58% higher compared to 14.30 g/s, the maximum amount of vapor emission when S was 20 L/min. For each case, i.e., S = 60 L/min and S = 20 L/min, the corresponding maximum condenser duty for the total condensation was −61.56 and −39.37 kW, respectively. Thus, in this case, the smaller the capacity of the vacuum pump, the better the process design. Moreover, if the capacity was too large, sudden boiling (socalled bumping phenomenon) was expected to occur. It was therefore advantageous to decrease the pressure as slowly as possible. However, the capacity of the vacuum pump could not be reduced unconditionally due to the limitation on the duration of vacuum operation as well as the pump’s capacity itself. The pressure should reach the target set point within the total process duration of 1 h. Therefore, as shown in Figure 9, at least 20 L/min of the pump would need to be used to pull the vacuum from the free volume of 200 L. Conversely, the result was different for the mixing ratio of 7:1. As shown in Figure 8b, the quantity of vapor generated

Table 1. Simulation Results of Basic Design

total process time (s) depressurizing interval (s) isobaric interval (s) maximum vapor flow (g/s) maximum cooling duty (kW) initial feed temperature (°C) minimum temperature (°C) final feed temperature (°C) initial holdup (kg) final holdup (kg) initial composition (wt % H2SO4) final composition (wt % H2SO4)

value

occurrence time (s)

3395 767 2628 16.2 −44.63 163 137 182 160 137 80 93

− − − 220 220 0 767 3395 0 3395 0 3395

(8)

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duty for the total condensation was −10.01 and −18.66 kW, respectively. Thus, the higher the capacity of the vacuum pump, the better the process design. This result was opposite of the outcome discussed earlier herein. The detailed simulation results of the above two cases are represented in Table S3 in the Supporting Information. Figure 10 shows the maximum vapor flows for all the feed mixtures at different capacities of the vacuum pump. The

Figure 10. Maximum vapor generated at different capacities of vacuum pump and different feed conditions.

mixtures were divided into two groups. For the feeds with mixing ratios of 2:1, 3:1, and 4:1, the amount of vapor generated continued to increase as the capacity of the vacuum pump increased. However, for the feeds with mixing ratios of 5:1, 6:1, and 7:1, the value increased to a certain extent and then decreased again, forming peaks. The peaks formed because of the bubble point shifting effect (see Figure 11). The bubble point changed at each moment not only by the pressure effect but also the temperature effect owing to the simultaneous occurrence of heating and vacuum operations. When the feed liquid being warmed up met the bubble point during the depressurization step (Figure 11c), more vapor was generated as the capacity of the vacuum pump increased. This was similar to the three simulation results in the case of the mixing ratio of 7:1 in Figure 8b and Figure 10, at S = 20, 30, and 40 L/min. However, when the pressure was reduced rapidly by the highcapacity vacuum pump, the feeds did not have a chance to attain the bubble point during the depressurization phase (Figure 11d), and, hence, they boiled in the isobaric phase. As mentioned before, the amount of instantaneous vapor generation in the isobaric period was extremely low compared to that in the depressurization step, similar to the other simulation results for the mixing ratio of 7:1 in Figure 10, at S = 50−100 L/min. Therefore, it was advantageous to use the lowcapacity vacuum pump for the first group, and the converse for the second one. The rationale behind adopting the converse operation strategies for each feed condition could be explained by Figure 12, wherein the initial feed temperature, boiling temperature at atmospheric pressure, and target low pressure are represented for each feed condition. As mentioned in section 2.1, the feed temperature decreased as the volumetric mixing ratio of sulfuric acid to hydrogen peroxide increased. This was because the

Figure 8. Instantaneous vapor emission for mixing ratios of (a) 2:1, and (b) 7:1 at various capacities of vacuum pump.

Figure 9. Vacuum operation times at various capacities of vacuum pump.

increased as S increased from 20 to 40 L/min. For capacities over 50 L/min, however, the instantaneous vapor flow in the system was greatly reduced and even the peak did not appear. Notably, for 100 L/min, the trend was almost the same as that for 50 L/min. When S was 100 and 20 L/min, the corresponding maximum amount of vapor was calculated as 3.74 and 6.88 g/s, respectively, and the maximum condenser 8549



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Figure 11. Bubble point shift effect on the various operating condition: (a) 2:1 mixing ratio when vacuum pump capacity is 20 L/min; (b) 2:1 mixing ratio when vacuum pump capacity is 60 L/min; (c) 7:1 mixing ratio when vacuum pump capacity is 20 L/min; (d) 7:1 mixing ratio when vacuum pump capacity is 100 L/min.

mbar, while for the volumetric mixing ratios of 5, 6, and 7, the feed temperature was still below the boiling point at 42.5 mbar. From the viewpoint of VLE, the former mixtures were now superheated vapor at 42.5 mbar. In other words, if the pressure changed suddenly from 1013 mbar to 42.5 mbar, the former mixtures abruptly changed the state from subcooled liquid to vapor. In terms of the vacuum distillation process, when high capacity of the vacuum pump was used, a huge amount of vapor was generated in the tank and, consequently, high condensing duty was imposed on the condenser. On fixing the heating rate and increasing the vacuum pump capacity, the rate of evaporation increases significantly, as illustrated by the difference in temperature gradients; see Figure 11a and b. On the contrary, the mixing ratios of 5, 6, and 7 maintained the initial state of the subcooled liquid. These mixtures still remained in the liquid phase whose vapor fraction was zero, even though the pressure changed from 1013 mbar to 42.5 mbar. Thus, in this case, no matter how rapidly the pressure was reduced, the mixtures did not transform from the original liquid state to vapor. Therefore, by pulling the vacuum as fast as possible, the sudden bumping phenomena similar to the cases of S = 20, 30, and 40 L/min in Figure 8b could be prevented before the boiling point shift because of the heating. 4.3. Analysis of Effects on Alternative Operating Condition. In section 2.3, the operating condition for the base case process was chosen from the phase diagram. In terms of cooling, the target pressure was determined from the VLE of

Figure 12. Comparison of boiling points at 1013 mbar and 42.5 mbar with feed temperatures at various feed conditions.

hydrogen peroxide, a limiting reactant which had participated in the exothermic reaction, was reduced. For all the cases, the feed temperature was below the normal boiling point; i.e., initially, the feed was in the subcooled liquid state. However, with decreasing pressure, the boiling point decreased, and a reversal of events occurred. For the volumetric mixing ratios of 2, 3, and 4, the feed temperature was above the boiling point at 42.5 8550



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water, and then the bubble point of 93 wt % H2SO4 solution at that pressure was chosen as the target temperature. However, for certain reasons, it was likely that the operating condition would need to be revised and, consequently, the process design would be accordingly modified. First, the operating pressure should be increased if there are severe restrictions on the cooling. For example, if the cooling temperature of 40 °C was given and if we assumed 10 °C as the minimum temperature difference over the heat exchanger, the minimum condensing temperature would now be computed as 50 °C. From Figure 5, the operating pressure could not be set below 123.5 mbar, the dew point of the pure steam at 50 °C. This was because if the pressure was lower than the aforementioned value, the total condensation could not be guaranteed. Alternately, it may be necessary for the operating temperature to not exceed a certain value if an upper limit of the process temperature was given. To illustrate, when the construction materials are limited, either environmental issues emerge or the process itself restricts the temperature. For example, if we took 150 °C as the target operating temperature, the Pxy phase diagram (Figure 6) at that temperature revealed that at 10.9 mbar, 93 wt % H2SO4 liquid was saturated with the vapor. The condensation temperature at such low pressure was 8.24 °C. In this last section, we will see how the strategies for vacuum operation should be changed for the two cases above, i.e., operating pressures of 123.5 and 10.9 mbar. The results are summarized in Table 2. For each situation, the key advantage

Figure 13. Comparison of boiling points at 10.9, 42.5, and 123.5 mbar with feed temperatures at various feed conditions.

when the feed was below its bubble point. However, given that the boiling point essentially is a function of the pressure, the boiling curve shifted as the operating pressure changed. Compared with a basis of 43.5 mbar, the curve shifted upward when the pressure increased to 123.5 mbar. In this case, the feed with the mixing ratio of 4:1 changed its state at the target pressure from vapor to subcooled liquid, and consequently the capacity of the vacuum pump strategically changed from “low” to “high”. Conversely, the curve shifted downward when the pressure decreased to 10.9 mbar. The feed with the mixing ratio of 5:1 changed its phase from subcooled liquid to vapor. Consequently, the capacity of the vacuum pump changed from high to low. Although the other feeds remained the same here, it was possible to change their vacuum operating strategies if the target pressure changed further. Thus, it was in principle possible to derive suitable vacuum strategies for all the cases discussed hereinabove by referring to the phase diagram.

Table 2. Capacities of Vacuum Pump for Various Mixing Conditions volumetric mixing ratios of sulfuric acid and hydrogen peroxide operating pressure (mbar)

2:1

3:1

4:1

5:1

6:1

7:1

10.9 42.5 123.5

low low low

low low low

low low high

low high high

high high high

high high high

5. CONCLUSIONS In this work, the design of the processes for the regeneration of waste sulfuric acid in the semiconductor industry was presented. A dynamic model of BVD was developed, followed by several sensitivity analyses on the feed condition, vacuum pump capacity, and cooling utility. By applying the main characteristics of the BVD processes discussed in section 4.1, we formulated feasible designs for the different conditions described in sections 4.2 and 4.3. For each feed condition, the capacity of the vacuum pump was selected such that the size of the condenser was minimized. As a result of model computation, it is concluded that at the lower mixing ratio of feed solution, it is desirable to lower the vacuum suction rate and vice versa. We also studied the methods to determine the operating conditions, and simulation of the batch distillation process involving vacuum pump through repetitive steady-state simulations. Although the solver does not give the exact results of BVD dynamics, the simulation method presented in this study is simple and easy to implement. The methodologies presented herein with regard to defining the design problem followed by approaching and solving it could be applied to other chemicals as well. There are several opportunities available to optimize the BVD technique. In the future, we plan to study control,

was the capacity of the vacuum pump. The expression “high” means that a high capacity is needed and vice versa. For comparison, the cases with 42.5 mbar treated in the previous section are given in the table. As mentioned before, it was advantageous to use the vacuum pump with low capacity for the mixing ratios of 2:1, 3:1, and 4:1. For the cases with 10.9 mbar, however, it was desirable to reduce the pressure as slowly as possible for mixing ratios of 2:1, 3:1, 4:1, and 5:1. Finally, for the cases with high operating pressure, 123.5 mbar, the low capacity of vacuum pump was used for only two mixing ratios, 2:1 and 3:1. Notably, for the feed with the mixing ratios of 4:1 and 5:1, the vacuum strategies varied as the target pressure changed. Moreover, the strategies did not change for the rest of the feed mixtures. The rationale behind using different capacities of the vacuum pump with the change in the operating pressure for certain feeds could also be explained by the phase diagram. Figure 13 represents the bubble temperature curves of the sulfuric acid solution at different pressures. The initial feed temperature is also shown in the plot. As stated in the previous section, it was advantageous to pull the vacuum as slowly as possible when the feed was above its bubble point at target pressure. On the other hand, it was preferable to reduce the pressure as fast as possible 8551



Article

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scheduling, and multicomponent separation to investigate the feasibility and sustainability of the process.

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ASSOCIATED CONTENT

S Supporting Information *

Calculation results for the heat emission of two reactions, i.e., decomposition reaction of hydrogen peroxide and dilution of sulfuric acid (Table S1), the feed condition (Table S2), and the simulation results of mixing ratios of 2:1 and 7:1 (Table S3). This material is available free of charge via the Internet at http://pubs.acs.org.

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AUTHOR INFORMATION

Corresponding Author

*Tel.: 82-2-880-1887. E-mail: [email protected]. Notes

The authors declare no competing financial interest.

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ACKNOWLEDGMENTS This research was supported by the second phase of the Brain Korea 21 Program in 2014, by the Institute of Chemical Processes in Seoul National University, by MKE and grant from the LNG Plant R&D Center funded by the Ministry of Land, Transportation, and Maritime Affairs (MLTM) of the Korean government, and by the Energy Efficiency & Resources Core Technology Program of the Korea Institute of Energy Technology Evaluation and Planning (KETEP) granted financial resource from the Ministry of Trade, Industry & Energy, Republic of Korea (No. 2010201020006D), (No. 20132010201760) and (No. 20132010500050).

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NOMENCLATURE ΔrH°298 [=] standard enthalpy of reaction (J/mol) αA,B [=] relative volatility ΔTmin [=] minimum temperature difference (K) T [=] temperature (K) x [=] liquid composition (dimensionless) y [=] vapor composition (dimensionless) S [=] vacuum pump’s capacity (L/min) αA,B [=] relative volatility (dimensionless) p [=] pressure (Pa) p0 [=] initial pressure (Pa) t [=] time (minute) V [=] free volume (L)

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REFERENCES

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Systematic Regeneration of Waste Sulfuric Acid in Semiconductor

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