Sulfur recovery method based on organic base catalyzed hydrolysis and spiral pressure filtration

Abstract

The present application relates to the technical field of industrial gas purification and sulfur resource recovery, and discloses a sulfur recovery method based on organic alkali catalytic hydrolysis and spiral pressure filtration, which comprises the following steps: feeding raw gas containing COS and / or CS2 and SO2 into an absorption reaction liquid containing organic alkali, aromatic hydrocarbon solvent and non-ionic surfactant, and reacting at 120-150 DEG C to generate liquid sulfur; cooling the reaction liquid to below the sulfur freezing point to obtain a sulfur-containing slurry; performing solid-liquid separation by a spiral pressure filter to obtain sulfur cake and filtrate; heating the filtrate to above the cloud point temperature and standing to separate into layers, and recycling the upper layer regeneration liquid. Through the multifunctional synergistic effect of the non-ionic surfactant in dispersing liquid sulfur in the reaction section, regulating the particle morphology in the crystallization section and enriching impurities in the regeneration section, combined with the spiral pressure filtration mechanical separation and cloud point regeneration technology, the present application realizes efficient removal of various sulfides and recovery of sulfur resources, effectively solves the problem of sulfur plugging, and guarantees long-period stable operation of the system.

Description

Technical Field

[0001] This invention relates to the field of industrial gas purification and sulfur resource recovery technology, and in particular to a sulfur recovery method based on organic base catalytic hydrolysis and spiral pressure filtration. Background Technology

[0002] In the processes of coal chemical industry, natural gas purification, oil refining, and syngas production, the industrial tail gas or raw gas often contains sulfur-containing impurities such as carbonyl sulfide (COS), carbon disulfide (CS2), and sulfur dioxide (SO2). The presence of these sulfides not only causes equipment corrosion and poisoning of catalysts in subsequent processes, but their emissions also lead to serious environmental pollution problems.

[0003] One of the mainstream technologies for treating sulfur-containing gases in industry is the modified Claus process, which can convert hydrogen sulfide into elemental sulfur for recovery. However, when the feed gas contains COS and CS2, the traditional Claus process faces significant limitations. COS and CS2 are chemically stable and difficult to convert directly under conventional Claus reaction conditions; they typically require pre-conversion to H2S via hydrolysis before entering the Claus system. In existing technologies, the hydrolysis of COS and CS2 is mostly carried out using solid catalysts in fixed-bed reactors, with reaction temperatures usually controlled between 60 and 100°C. This approach presents several insurmountable problems. First, the hydrolysis reaction and the subsequent Claus reaction need to be carried out in different temperature zones, resulting in a longer process flow and increased equipment investment. Second, the hydrolysis reaction temperature is below the melting point of elemental sulfur, which may generate solid sulfur during the reaction. This solid sulfur easily deposits in the catalyst pores and the inner walls of the reaction pipes, leading to decreased catalyst activity and increased system pressure drop, and in severe cases, causing blockage and forced shutdown. Furthermore, solid catalysts are prone to sulfation poisoning in aqueous and oxygen-containing atmospheres, resulting in a short service life and requiring frequent replacement.

[0004] Another approach is to perform sulfide conversion in the liquid phase. Related studies have attempted to treat sulfur-containing gases in acidic aqueous media, utilizing the acidic environment to promote the Claus reaction. However, acidic conditions significantly inhibit the hydrolysis of COS and CS2, resulting in low removal efficiency for these two key sulfides. Although some studies have proposed introducing organic bases in subsequent steps to promote hydrolysis, this segmented approach disrupts the reaction system, failing to achieve synergistic removal of multiple sulfides, and also faces the problem of sulfur product deposition and blockage in practical operation.

[0005] In summary, existing technologies for treating industrial exhaust gases containing COS and CS2, whether employing solid catalyst hydrolysis or acidic media treatment, have failed to effectively address the core challenge of sulfur product clogging of equipment. Furthermore, they struggle to achieve efficient synergistic removal of multiple sulfides and comprehensive sulfur resource recovery within the same reaction system. Therefore, developing a method that fundamentally avoids sulfur clogging and achieves integrated removal of multiple sulfides and sulfur recovery is of significant practical importance for industrial exhaust gas purification and sulfur resource utilization. Summary of the Invention

[0006] The technical problem solved by this invention is that the existing technology has problems such as complex process flow, easy clogging of equipment by sulfur products, inability to achieve synergistic and efficient removal of multiple sulfides, and low sulfur resource recovery efficiency.

[0007] To address the above problems, the present invention provides the following technical solution:

[0008] This invention provides a sulfur recovery method based on organic base catalytic hydrolysis and spiral pressure filtration, comprising the following steps:

[0009] The feed gas containing COS and / or CS2 and SO2 gas are passed into an absorption reaction tower containing an absorption reaction liquid and reacted at 120-150°C. The COS and / or CS2 are hydrolyzed to generate H2S, and the generated H2S reacts with SO2 in a liquid-phase Claus reaction to generate elemental sulfur. The absorption reaction liquid contains an organic base, an aromatic solvent, a nonionic surfactant, and water.

[0010] The absorbent reaction liquid containing liquid sulfur obtained after the reaction is cooled to below the freezing point of sulfur to obtain a slurry containing sulfur solid particles.

[0011] The slurry is fed into a screw filter press for solid-liquid separation to obtain sulfur filter cake and filtrate.

[0012] The filtrate is heated to above the cloud point temperature of the nonionic surfactant, allowed to stand and separate into layers, and the upper layer of regenerated organic alkaline aqueous solution and the lower layer of heavy phase rich in surfactant and impurities are obtained.

[0013] The regenerated organic alkaline aqueous solution in the upper layer is returned to the absorption reaction tower for recycling.

[0014] By adopting the above technical solution, the present invention achieves the synergistic removal of COS, CS2 and SO2 and sulfur recovery in the same reaction system, solves the sulfur blockage problem, and realizes the recycling and regeneration of the absorption reaction liquid.

[0015] Specifically, within the absorption reaction tower, the organic base plays a dual catalytic role. The nitrogen atom in the organic base molecule possesses a lone pair of electrons, which can act as a nucleophile to attack the carbon atom in the COS or CS2 molecules, forming an intermediate complex. This lowers the activation energy of the hydrolysis reaction, promoting the hydrolysis of COS and CS2 to produce H2S and CO2. The reaction equations are as follows:

[0016] COS + H2O → H2S + CO2;

[0017] CS2+ 2H2O → 2H2S + CO2;

[0018] Meanwhile, the weak acid salts formed by organic bases and H2S can act as reservoirs for H2S, increasing the concentration of H2S in the liquid phase. - The concentration of H2S increases, thus accelerating the liquid-phase Claus reaction between H2S and SO2 to produce elemental sulfur. The reaction equation is as follows:

[0019] 2H2S + SO2→ 3 / x Sx + 2H2O;

[0020] The reaction temperature is controlled at 120–150°C, higher than the melting point of elemental sulfur (approximately 115°C). Therefore, the generated elemental sulfur is dispersed in the absorption reaction solution as liquid droplets. During this reaction stage, although the system temperature is higher than the conventional cloud point of nonionic surfactants, and the surfactants become sharply more hydrophobic due to high-temperature dehydration, the continuously generated, highly hydrophobic liquid sulfur droplets, in a dynamic gas-liquid mixing state, allow surfactant molecules to directionally adsorb their hydrophobic groups onto the surface of the liquid sulfur droplets. Simultaneously, the high concentration of organic base in the absorption reaction solution alters the polarity of the aqueous solvent system, maintaining good solvation of the hydrophilic chains of the surfactants at high temperatures. This allows the nonionic surfactants in the absorption reaction solution to adsorb onto the surface of the liquid sulfur droplets through their hydrophobic groups in the reaction section. The hydrophilic chains extend into the aqueous phase, forming a steric hindrance layer that prevents the sulfur droplets from colliding and coalescing, thus avoiding their sedimentation and adhesion within the reaction tower.

[0021] Subsequently, the sulfur-rich absorption reaction solution is cooled below the sulfur freezing point (50–70°C), causing the liquid sulfur to solidify into sulfur solid particles. During this crystallization stage, the selective adsorption of nonionic surfactant molecules on the sulfur crystal surface alters the growth rate of each crystal facet, inhibiting anisotropic crystal growth and resulting in uniformly sized, smooth particles, thus avoiding filtration difficulties caused by needle-like or flaky crystals. The resulting sulfur-containing solid particle slurry, after cooling, enters a screw filter press for solid-liquid separation. Mechanical extrusion rapidly separates the sulfur filter cake and filtrate, significantly shortening the separation time and reducing the moisture content of the filter cake.

[0022] The separated filtrate is heated to above the cloud point temperature of the nonionic surfactant (90–98°C). In this regeneration section, since the filtrate has undergone the aforementioned solid-liquid separation, most of the elemental sulfur in the system has been removed, and the surfactant loses the strongly hydrophobic interface upon which it relies for adsorption. Under conditions of lack of a hydrophobic interface and in a static environment, the solubility of the nonionic surfactant in water decreases with increasing temperature. When the temperature reaches the cloud point, the surfactant molecules change from a monomolecular dispersion state to a micelle aggregate state, precipitating from the aqueous phase to form an independent liquid phase. This phase transition process simultaneously carries away the hydrophobic impurities adsorbed on its surface, including sulfur oligomers, heavy aromatics, and organic alkali degradation products, forming a heavy phase rich in surfactants and impurities that settles at the bottom of the tank. The upper layer is a regenerated organic alkali aqueous solution, which can be returned to the absorption reaction tower for recycling. This process achieves continuous removal of impurities and in-situ regeneration of the absorption reaction liquid, effectively preventing solvent degradation caused by impurity accumulation.

[0023] This invention, through the aforementioned technical solution, constructs a V-shaped temperature gradient consisting of a high-temperature reaction section (120–150°C), a low-temperature crystallization section (50–70°C), and a medium-temperature regeneration section (90–98°C). The high-temperature reaction ensures that sulfur exists in a liquid state, preventing equipment blockage; the low-temperature crystallization solidifies the sulfur into solid particles, facilitating mechanical separation; and the medium-temperature regeneration utilizes the cloud point characteristic to allow surfactant precipitation, avoiding the damage to the surfactant caused by high temperatures. This temperature gradient design enables sulfur to complete a cycle of liquid generation, solid-state separation, and liquid regeneration within the system, completely eliminating the risk of sulfur blockage.

[0024] Preferably, in the absorption reaction solution, the organic base accounts for 30%–50% by mass, the aromatic solvent accounts for 5%–10% by mass, the nonionic surfactant accounts for 0.5%–3.0% by mass, and the balance is water. Within this concentration range, the organic base can provide sufficient catalytic active centers while avoiding increased costs due to excessively high concentrations and decreased reaction efficiency due to excessively low concentrations. The aromatic solvent, as a co-solvent for liquid sulfur, can reduce the viscosity of sulfur and promote its dispersion in the reaction solution. The nonionic surfactant, within the range of 0.5%–3.0%, can effectively perform its functions of dispersion, regulation, and enrichment without unnecessary consumption due to excessively high concentrations.

[0025] Preferably, the reaction temperature inside the absorption reaction tower is 130–150°C; the cooling temperature in the cooling step is 50–70°C; and the heating temperature in the filtrate heating step is 90–98°C. The reaction temperature of 130–150°C ensures that sulfur exists in a liquid state while providing suitable hydrolysis reaction kinetics. The cooling temperature of 50–70°C ensures complete solidification of liquid sulfur without increasing energy consumption due to excessive cooling. The heating temperature of 90–98°C falls precisely within the cloud point range of nonionic surfactants, achieving sufficient cloud point separation while avoiding degradation of surfactants and organic bases caused by excessively high temperatures.

[0026] Preferably, the molar ratio of the SO2 gas introduced to the H2S generated after the complete hydrolysis of COS and / or CS2 in the feed gas is 1:1.8 to 1:2.2. This ratio is close to the theoretical stoichiometric ratio of the Claus reaction (1:2), ensuring that H2S and SO2 react fully and avoiding excess of one reactant leading to unreacted sulfides remaining in the tail gas.

[0027] Preferably, the feed pressure of the screw filter press is 0.3–0.8 MPa, and the screw compression ratio is 2:1–4:1. During the filtration process, a cooling jacket maintains the internal temperature below the melting point of sulfur. Appropriate feed pressure and compression ratio ensure efficient solid-liquid separation while preventing excessive compression that could lead to an overly dense filter cake and affect the washing effect. The cooling jacket ensures that the sulfur remains solid during filtration, preventing it from remelting due to frictional heat and adhering to the equipment.

[0028] Preferably, the sulfur filter cake separated by the screw filter press is reverse-washed with an aromatic solvent at the end of the filter press, and then dried to obtain the sulfur product. Reverse washing can effectively remove residual organic alkalis and surfactants from the filter cake, improving the purity of the sulfur product.

[0029] Preferably, the COS content in the feed gas is 0.1%–5%, and the CS2 content is 0.05%–3%. The method of the present invention can maintain stable desulfurization efficiency and sulfur product quality within this range of feed gas composition, demonstrating good adaptability to feed gas.

[0030] Preferably, the organic base is pyridine, indole, or 8-hydroxyquinoline. These three aromatic heterocyclic organic bases have nitrogen atoms in a conjugated system, resulting in high electron cloud density, strong nucleophilicity, and excellent catalytic activity for the hydrolysis of COS and CS2. The aromatic solvent is trimethylbenzene or a heavy aromatic hydrocarbon, which has good solubility for liquid sulfur. The nonionic surfactant is an alkylphenol polyoxyethylene ether with a cloud point of 60–78°C and an HLB value of 12–14. This type of surfactant can exert multifunctional synergistic effects of dispersion, regulation, and enrichment in the reaction, crystallization, and regeneration stages, respectively, and is a key component in achieving the technical solution of this invention.

[0031] In summary, the present invention has at least one of the following beneficial technical effects:

[0032] 1. This invention controls the reaction temperature above the sulfur melting point, so that the generated sulfur exists in liquid form, avoiding the solidification and blockage of solid sulfur in the equipment; at the same time, the dispersing effect of nonionic surfactants prevents the agglomeration and deposition of liquid sulfur droplets, fundamentally solving the sulfur blockage problem in traditional sulfur recovery processes and achieving long-term stable operation of the system.

[0033] 2. This invention utilizes the dual catalytic effect of organic bases to simultaneously achieve the hydrolysis conversion of COS and CS2 as well as the liquid-phase Claus reaction of H2S and SO2 within the same reaction system, avoiding the complexity of multiple processes in series.

[0034] 3. This invention achieves an integrated design for liquid sulfur dispersion, sulfur particle morphology regulation, and impurity enrichment and removal through the multifunctional synergistic effect of nonionic surfactants in the reaction, crystallization, and regeneration stages, significantly simplifying the process flow.

[0035] 4. This invention utilizes the cloud point characteristics of nonionic surfactants to achieve in-situ regeneration of solvent and continuous removal of impurities in a heated clarification tank, effectively preventing solvent deterioration caused by impurity accumulation.

[0036] 5. The sulfur product produced by this invention has a purity of over 99.5%, and its ash content, acidity, moisture content, and organic matter content are all superior to the GB / T 2449-2006 standard for first-grade industrial sulfur, thus achieving high-quality recovery of sulfur resources.

[0037] 6. The method of the present invention has wide adaptability to the composition of raw gas, can treat various industrial tail gases, and is applicable to multiple fields such as coal chemical industry, natural gas purification, and oil refining. Attached Figure Description

[0038] Figure 1 This is a schematic diagram of the sulfur recovery device based on organic base catalytic hydrolysis and spiral pressure filtration according to the present invention.

[0039] Figure 2 The diagram shows a comparison of the catalytic performance of different organic base systems of the present invention, where (a) is a comparison of the conversion rates of COS and CS2, and (b) is a comparison of the H2S generation rate and the liquid-phase Claus reaction rate.

[0040] Figure 3 The graphs show the changes in the particle size of liquid sulfur microdroplets over time for different surfactant systems of the present invention, where (a) is a comparison graph of the particle size changes over time, and (b) is a comparison graph of the initial D50 and the D50 after 60 min.

[0041] Figure 4 The following are comparative graphs showing the sulfur particle filtration performance of different embodiments and comparative examples of the present invention; wherein (a) is a comparison graph of particle size distribution, and (b) is a comparison graph of filter cake moisture content, filtration time, and filter press operating torque.

[0042] Figure 5 The following are graphs showing the solution stability test results of different systems of the present invention; where (a) is a comparison graph of the changes in organic base content and surfactant content over time, and (b) is a comparison graph of the changes in TOC and total impurity content over time.

[0043] The components include: 1. Absorption reaction tower; 2. Cooler; 3. Screw filter press; 4. Heating clarification tank; 5. Circulating pump; 6. Filter; and 7. Sulfur collection device. Detailed Implementation

[0044] To make the objectives, technical solutions, and advantages of this invention clearer, the invention will be further described in detail below with reference to specific embodiments. It should be understood that the specific embodiments described herein are merely illustrative and not intended to limit the invention. Any modifications, equivalent substitutions, or improvements made by those skilled in the art without departing from the spirit and scope of this invention should be included within the protection scope of this invention.

[0045] Please see the appendix Figure 1 This invention provides a sulfur recovery device based on organic base catalytic hydrolysis and spiral pressure filtration, the device comprising:

[0046] Absorption reaction tower 1 has a raw material gas inlet and a sulfur dioxide inlet at the bottom, a purified gas outlet and an absorbent liquid inlet at the top, and a reaction liquid outlet at the bottom. Absorption reaction tower 1 is used to react raw material gas containing COS and / or CS2 with sulfur dioxide in an organic alkaline aqueous solution to generate a mixture containing liquid sulfur.

[0047] Cooler 2, whose inlet is connected to the reaction liquid outlet at the bottom of absorption reaction tower 1, is used to cool the reaction liquid below the sulfur freezing point, so that the liquid sulfur solidifies to form a slurry containing sulfur solid particles.

[0048] The screw filter press 3, whose inlet is connected to the outlet of the cooler 2, is used for solid-liquid separation of slurry containing sulfur solid particles to obtain sulfur filter cake and filtrate. The screw filter press 3 is provided with a filter cake outlet and a filtrate outlet, and the outer shell is provided with a cooling jacket (not shown in the figure) to maintain the internal temperature below the melting point of sulfur.

[0049] The heating clarification tank 4, whose inlet is connected to the filtrate outlet of the screw filter press 3, is used to heat the filtrate and allow it to settle and separate into layers. Utilizing the cloud point characteristics of surfactants, the surfactant-rich phase precipitates and settles. The upper layer is a regenerated organic alkaline aqueous solution, and the lower layer is a heavy phase rich in surfactants and impurities. The heating clarification tank 4 is equipped with an upper clear liquid outlet and a bottom slag discharge outlet.

[0050] The circulating pump 5 has its inlet connected to the upper clear liquid outlet of the heating and clarifying tank 4 and its outlet connected to the absorbent inlet of the absorption reaction tower 1. It is used to send the regenerated organic alkali aqueous solution back to the absorption reaction tower for recycling.

[0051] Filter 6 is installed on the pipeline between circulating pump 5 and absorption reaction tower 1 to filter out trace solid impurities that may be entrained in the regeneration solution.

[0052] The sulfur collection device 7 is connected to the filter cake outlet of the screw filter press 3 and is used to collect sulfur filter cake, which is then dried to obtain sulfur products.

[0053] Optionally, the device further includes a heat exchanger (not shown) for precooling the reaction liquid before it enters the cooler 2, or for preheating the filtrate before it enters the heating clarification tank 4.

[0054] The absorption reaction tower 1 can be a bubbling tower, a spray tower, a packed tower, or a composite tower combining bubbling and spraying. Internal components such as packing layers or trays can be installed inside the tower to enhance gas-liquid contact. A heating and insulation jacket or insulation layer can be installed on the outer wall of the tower to maintain the reaction temperature.

[0055] The spiral filter press 3 is a conventional spiral extrusion solid-liquid separation device. Its spiral compression ratio, feed pressure, and spiral speed can be adjusted according to the material characteristics. Cooling medium is introduced into the cooling jacket to ensure that sulfur remains solid during the filtration process.

[0056] The heated clarification tank 4 is a settling tank with heating function. It can maintain the temperature inside the tank by using steam coils or electric heating, and is equipped with a sight glass and a level gauge to observe the stratification interface.

[0057] This device, through the organic combination of its various units, achieves continuous operation of high-temperature reaction, cooling crystallization, mechanical pressure filtration, and solvent regeneration, ensuring efficient sulfur recovery and long-term stable operation of the system.

[0058] See appendix Figure 1 The method for sulfur recovery using the above-mentioned device includes the following steps:

[0059] First, a feed gas containing COS and / or CS2 is introduced into the lower part of absorption reactor 1 through the feed gas inlet, while sulfur dioxide gas is simultaneously introduced into the lower part of absorption reactor 1 through the sulfur dioxide inlet. The absorption reaction liquid (an aqueous solution of an organic alkali) is sprayed down from the top of the tower through the absorption liquid inlet, contacting the rising gas flow counter-currently. Inside absorption reactor 1, the reaction temperature is controlled above the melting point of sulfur (120–150°C). Under the catalytic action of the organic alkali, COS and / or CS2 undergo hydrolysis to produce H2S and CO2. The generated H2S then undergoes a liquid-phase Claus reaction with SO2 to produce elemental sulfur. Because the reaction temperature is above the melting point of sulfur, the generated elemental sulfur is dispersed in the absorption reaction liquid in the form of liquid droplets.

[0060] Subsequently, the absorption reaction liquid rich in liquid sulfur is discharged from the reaction liquid outlet at the bottom of the absorption reaction tower 1 and enters the cooler 2. In the cooler 2, the reaction liquid is cooled to below the freezing point of sulfur (50-70°C), and the liquid sulfur solidifies to form sulfur solid particles, resulting in a slurry containing sulfur solid particles.

[0061] The slurry containing sulfur solid particles enters the screw filter press 3 for solid-liquid separation. During the filtration process, a cooling medium is introduced through the cooling jacket of the outer shell to maintain the internal temperature below the melting point of sulfur, ensuring that the sulfur remains solid. The sulfur filter cake obtained from the filtration is discharged through the filter cake outlet and enters the sulfur collection device 7, where it is dried to obtain the sulfur product; the filtrate obtained from the filtration is discharged through the filtrate outlet and enters the heating clarification tank 4.

[0062] The filtrate is heated to 90-98°C in the heating and clarification tank 4. The cloud point characteristics of the nonionic surfactant cause it to precipitate from the aqueous phase, forming a heavy phase rich in surfactant and impurities that settles to the bottom of the tank and is periodically discharged from the bottom slag outlet. The upper layer is a regenerated organic alkali aqueous solution, which flows out through the upper clear liquid outlet, is transported by the circulation pump 5, and after being filtered by the filter 6 to remove any trace solid impurities that may be entrained, it is returned to the absorption liquid inlet at the top of the absorption reaction tower 1 for recycling.

[0063] The above steps are carried out continuously to achieve the purification of sulfur-containing raw gas, sulfur recovery, and regeneration and recycling of the absorption reaction liquid.

[0064] The present invention will be further described in detail below through specific embodiments.

[0065] The main raw materials and reagents used in the following examples and comparative examples have the following sources and specifications. Reagents not specifically mentioned are all commercially available analytical grade or higher grade products.

[0066] Pyridine (C5H5N, CAS No.: 110-86-1), molecular weight 79.10, melting point -41.6℃, boiling point 115.2~115.3℃, density 0.98g / mL, industrial grade purity ≥99.0%.

[0067] Indole (C8H7N, CAS No.: 120-72-9), molecular weight 117.15, melting point 52.5℃, boiling point 254℃, density 1.22g / mL, industrial grade purity ≥99.0%.

[0068] 8-Hydroxyquinoline (C9H7NO, CAS No.: 148-24-3), molecular weight 145.16, melting point 76℃, boiling point approximately 267℃, density 1.0015 g / mL, industrial grade purity ≥99.0%.

[0069] Trimethylbenzene (C9H) 12 (CAS No.: 25551-13-7), molecular weight 120.19, boiling range 164~176℃, density 0.86~0.88g / mL, industrial grade purity ≥98.0%.

[0070] Heavy aromatics, a mixture of C9 to C10 aromatics, mainly containing components such as trimethylbenzene, ethylbenzene, and propane, with a distillation range of 150 to 200°C, a density of 0.87 to 0.90 g / mL, and an industrial grade aromatic content of ≥95.0%.

[0071] TX-10 (nonylphenol polyoxyethylene ether-10) has an average molecular weight of approximately 660, a hydroxyl value of 80–90 mg KOH / g, a cloud point (1% aqueous solution) of 60–67 °C, an HLB value of 12–13, a moisture content of ≤1.0%, and an active ingredient content of ≥98.0%.

[0072] OP-10 (octylphenol polyoxyethylene ether-10) has an average molecular weight of approximately 646, a hydroxyl value of 83–91 mg KOH / g, a cloud point (1% aqueous solution) of 68–78 °C, an HLB value of 13.3–14.0, a moisture content of ≤1.0%, and an active ingredient content of ≥98.0%.

[0073] All other routine reagents not specifically mentioned are commercially available industrial-grade products.

[0074] Example 1:

[0075] This embodiment provides a sulfur recovery method based on organic base catalytic hydrolysis and spiral pressure filtration, including the following steps:

[0076] Weigh out 40% pyridine, 7% trimethylbenzene, 1.5% OP-10, and the remainder water by mass percentage, and stir and mix at 40°C for 45 minutes to prepare the absorption reaction solution.

[0077] The absorption reaction solution was applied at a flow rate of 2.5 m. 3 A feed gas containing 1.2% COS, 0.5% CS2, and 98.3% N2 is continuously fed to the top of the absorption reactor at a flow rate of / h, with the reaction temperature inside the reactor controlled at 140±2℃ and the gauge pressure at 1.0~1.5kPa.3 A flow rate of / h is introduced from the bottom of the tower, along with 10.7m³ of SO2 gas. 3 The SO2 feed rate is 1:2.05, and the molar ratio of SO2 to H2S generated after complete hydrolysis of COS and CS2 in the feed gas is 2.5 L / m³. 3 .

[0078] The reaction liquid discharged from the bottom of the absorption reaction tower is cooled to 60±2℃ at a rate of 10℃ / min through a heat exchanger to obtain a slurry containing sulfur solid particles. The slurry is continuously fed into a screw filter press, with the feed pressure controlled at 0.5MPa, screw compression ratio at 3:1, and screw speed at 5rpm. Cooling water at 25℃ is circulated through the cooling jacket of the filter press to maintain the internal temperature ≤65℃. Trimethylbenzene is sprayed into the washing section at the end of the filter press at a spray volume of 8% of the filter cake mass. The filter cake separated by the filter press is dried at 60-80℃ under normal pressure to obtain the sulfur product. The separated filtrate is sent to a heated clarification tank, with the temperature controlled at 95±2℃ and allowed to stand under normal pressure for 90 minutes. The lower surfactant-rich phase is discharged every 12 hours, and the upper aqueous phase is returned to the absorption reaction tower for recycling.

[0079] Example 2:

[0080] Weigh out 40% pyridine, 7% trimethylbenzene, 0.5% OP-10, and 52.5% water by mass percentage, and stir and mix them at 40°C for 45 minutes to prepare the absorption reaction solution.

[0081] SO2 gas was introduced at a rate of 10.7 m. 3 / h, other operating steps and process parameters are the same as in Example 1.

[0082] Example 3:

[0083] Weigh out 40% pyridine, 7% trimethylbenzene, 3.0% OP-10, and 50.0% water by mass percentage, and stir and mix at 40°C for 45 minutes to prepare the absorption reaction solution.

[0084] SO2 gas was introduced at a rate of 10.7 m. 3 / h, other operating steps and process parameters are the same as in Example 1.

[0085] Example 4:

[0086] Weigh out 40% pyridine, 5% trimethylbenzene, 1.5% OP-10, and 53.5% water by mass percentage, and stir and mix at 40°C for 45 minutes to prepare the absorption reaction solution.

[0087] SO2 gas was introduced at a rate of 10.7 m. 3 / h, other operating steps and process parameters are the same as in Example 1.

[0088] Example 5:

[0089] Weigh out 40% pyridine, 10% trimethylbenzene, 1.5% OP-10, and 48.5% water by mass percentage, and stir and mix them at 40°C for 45 minutes to prepare the absorption reaction solution.

[0090] SO2 gas was introduced at a rate of 10.7 m. 3 / h, other operating steps and process parameters are the same as in Example 1.

[0091] Example 6:

[0092] Weigh out 30% pyridine, 7% trimethylbenzene, 1.5% OP-10, and 61.5% water by mass percentage, and mix them at 40°C for 45 minutes to prepare the absorption reaction solution.

[0093] SO2 gas was introduced at a rate of 10.7 m. 3 / h, other operating steps and process parameters are the same as in Example 1.

[0094] Example 7:

[0095] Weigh out 50% pyridine, 7% trimethylbenzene, 1.5% OP-10, and 41.5% water by mass percentage, and stir and mix at 40°C for 45 minutes to prepare the absorption reaction solution.

[0096] SO2 gas was introduced at a rate of 10.7 m. 3 / h, other operating steps and process parameters are the same as in Example 1.

[0097] Example 8:

[0098] Weigh out 40% indole, 7% trimethylbenzene, 1.5% OP-10, and 51.5% water by mass percentage, and heat and stir at 65°C for 60 minutes until the indole is completely dissolved to prepare the absorption reaction solution.

[0099] SO2 gas was introduced at a rate of 10.7 m. 3 / h, other operating steps and process parameters are the same as in Example 1.

[0100] Example 9:

[0101] Weigh out 40% 8-hydroxyquinoline, 7% heavy aromatic hydrocarbons, 1.5% TX-10, and 51.5% water by mass percentage, and heat and stir at 85°C for 60 minutes until 8-hydroxyquinoline is completely dissolved to prepare the absorption reaction solution.

[0102] SO2 gas was introduced at a rate of 10.7 m. 3 / h, other operating steps and process parameters are the same as in Example 1.

[0103] Example 10:

[0104] Weigh out 40% pyridine, 7% trimethylbenzene, 1.5% OP-10, and 51.5% water by mass percentage, and stir and mix them at 40°C for 45 minutes to prepare the absorption reaction solution.

[0105] The reaction temperature inside the absorption tower was controlled at 130±2℃, and 10.7m³ of SO2 gas was introduced. 3 / h, other operating steps and process parameters are the same as in Example 1.

[0106] Example 11:

[0107] Weigh out 40% pyridine, 7% trimethylbenzene, 1.5% OP-10, and 51.5% water by mass percentage, and stir and mix them at 40°C for 45 minutes to prepare the absorption reaction solution.

[0108] The reaction temperature inside the absorption tower is controlled at 150±2℃. A condenser reflux device is installed at the top of the absorption tower to reduce the volatilization of organic base. 10.7m³ of SO2 gas is introduced. 3 / h, other operating steps and process parameters are the same as in Example 1.

[0109] Example 12:

[0110] Weigh out 40% pyridine, 7% trimethylbenzene, 1.5% OP-10, and 51.5% water by mass percentage, and stir and mix them at 40°C for 45 minutes to prepare the absorption reaction solution.

[0111] The reaction liquid discharged from the bottom of the absorption reaction tower is cooled to 50±2℃, and SO2 gas is introduced at a rate of 10.7m. 3 / h, other operating steps and process parameters are the same as in Example 1.

[0112] Example 13:

[0113] Weigh out 40% pyridine, 7% trimethylbenzene, 1.5% OP-10, and 51.5% water by mass percentage, and stir and mix them at 40°C for 45 minutes to prepare the absorption reaction solution.

[0114] The reaction liquid discharged from the bottom of the absorption reaction tower is cooled to 70±2℃, and 10.7m³ of SO2 gas is introduced. 3 / h, other operating steps and process parameters are the same as in Example 1.

[0115] Example 14:

[0116] Weigh out 40% pyridine, 7% trimethylbenzene, 1.5% OP-10, and 51.5% water by mass percentage, and stir and mix them at 40°C for 45 minutes to prepare the absorption reaction solution.

[0117] The feed gas composition is COS 0.5%, CS2 0.2%, N2 99.3%, and total sulfur content 0.7%, with a 1000m³ / h... 3 A flow rate of / h is fed into the absorption reactor, along with 4.7m³ of SO2 gas. 3 / h (the molar ratio of SO2 to H2S generated by hydrolysis is 1:2.05), and other operating steps and process parameters are the same as in Example 1.

[0118] Example 15:

[0119] Weigh out 40% pyridine, 7% trimethylbenzene, 1.5% OP-10, and 51.5% water by mass percentage, and stir and mix them at 40°C for 45 minutes to prepare the absorption reaction solution.

[0120] The feed gas composition is COS 2.5%, CS2 1.0%, N2 96.5%, and total sulfur content 3.5%, with a 1000m³ / h... 3 A flow rate of / h is fed into the absorption reactor, along with 22.5m³ of SO2 gas. 3 / h (the molar ratio of SO2 to H2S generated by hydrolysis is 1:2.00), and other operating steps and process parameters are the same as in Example 1.

[0121] Comparative Example 1:

[0122] Compared with Example 1, the difference is that OP-10 is not added to the absorption reaction solution. That is, the absorption reaction solution is prepared by weighing 40% pyridine, 7% trimethylbenzene, and 53% water by mass percentage. All other aspects are the same.

[0123] Comparative Example 2:

[0124] Compared with Example 1, the difference is that the organic base in the absorption reaction solution is replaced by triethylamine instead of pyridine. Specifically, the absorption reaction solution is prepared by weighing 40% triethylamine, 7% trimethylbenzene, 1.5% OP-10, and 51.5% water by mass percentage. All other aspects are the same.

[0125] Comparative Example 3:

[0126] Compared with Example 1, the difference is that thylene is not added to the absorption reaction solution. Instead, the absorption reaction solution is prepared by weighing 40% pyridine, 1.5% OP-10, and 58.5% water by mass percentage. All other aspects are the same.

[0127] Comparative Example 4:

[0128] Compared with Example 1, the difference is that the surfactant in the absorption reaction solution is sodium dodecylbenzenesulfonate instead of OP-10. That is, the absorption reaction solution is prepared by weighing 40% pyridine, 7% trimethylbenzene, 1.5% sodium dodecylbenzenesulfonate and 51.5% water by mass percentage. All other aspects are the same.

[0129] Comparative Example 5:

[0130] Compared with Example 1, the difference is that the reaction temperature inside the absorption reaction tower is controlled at 110±2℃, while the rest are the same.

[0131] Comparative Example 6:

[0132] Compared with Example 1, the difference is that the reaction liquid discharged from the bottom of the absorption reaction tower is cooled to 60±2℃ and then sent to a gravity settling tank for static stratification for 120 minutes instead of spiral pressure filtration separation. All other aspects are the same.

[0133] Comparative Example 7:

[0134] Compared with Example 1, the difference is that the reaction liquid (140±2℃) discharged from the bottom of the absorption reaction tower is directly sent to the screw filter press for pressure filtration and separation, without going through the cooling crystallization step, and the rest are the same.

[0135] Comparative Example 8:

[0136] Compared with Example 1, the difference is that the filtrate separated by the filter press is not treated by the heating clarification tank, but is directly returned to the absorption reaction tower for recycling. All other aspects are the same.

[0137] Comparative Example 9:

[0138] Compared with Example 1, the difference is that 20% pyridine, 7% trimethylbenzene, 1.5% OP-10 and 71.5% water were weighed according to the mass percentage to prepare the absorption reaction solution, and the rest were the same.

[0139] Comparative Example 10:

[0140] Compared with Example 1, the difference is that the temperature of the heating clarification tank is controlled at 80±2℃, while the rest are the same.

[0141] Test Example 1:

[0142] The experimental steps are as follows:

[0143] (1) Take 100 mL of each of the absorption reaction solution samples prepared in Examples 1 to 15 and place them in a stoppered conical flask, and label them for later use.

[0144] (2) The content of organic bases was determined by gas chromatography. Chromatographic conditions: HP-INNOWAX capillary column (30m×0.32mm×0.25μm), flame ionization detector, injection port temperature 250℃, detector temperature 280℃, column temperature programmed: initial temperature 80℃ held for 2min, then increased to 200℃ at 10℃ / min and held for 5min. The carrier gas was high-purity nitrogen, flow rate 1.5mL / min, split ratio 50:1. 1μL of the sample was directly injected after filtration through a 0.45μm filter membrane. Quantification was performed using the external standard method. The peak area of ​​the organic bases was recorded and the content was calculated. Each sample was measured in triplicate, and the average value was taken.

[0145] (3) The surfactant content was determined by ultraviolet spectrophotometry. Nonionic surfactants OP-10 and TX-10 have characteristic absorption in the ultraviolet region, with maximum absorption wavelengths of 275 nm and 277 nm, respectively. The sample was diluted with distilled water to an appropriate concentration, and the absorbance was measured at the corresponding wavelength using distilled water as a reference. The surfactant content was calculated based on the standard curve. The standard curve was plotted using OP-10 or TX-10 standard solutions of known concentrations. Each sample was measured in triplicate, and the average value was taken.

[0146] (4) The moisture content was determined by the Karl Fischer method. A micro moisture analyzer was used. 5 mL of the sample was injected into the titration cell and titrated to the endpoint with Karl Fischer reagent. The volume consumed was recorded and the moisture content was calculated. Each sample was measured in triplicate and the average value was taken.

[0147] (5) The cloud point temperature was determined by visual inspection. The sample solution (1% aqueous solution) was added to a test tube, a thermometer was inserted, and the tube was placed in a water bath and heated slowly while stirring. The temperature at which the solution changed from clear to cloudy was observed, which was the cloud point. Each sample was measured in parallel three times, and the average value was taken.

[0148] (6) The HLB value is calculated using the group number method; it is calculated by summing the number of hydrophilic and lipophilic groups based on the chemical structure of the surfactant. For OP-10 and TX-10, their HLB values ​​can also be estimated using the empirical relationship between cloud point and HLB, and then compared with the calculated results. In this test example, the HLB value is the average of the calculated value and the empirical value.

[0149] The experimental results are shown in Table 1:

[0150] Table 1. Performance test results of the absorption reaction solutions in Examples 1-15:

[0151]

[0152] According to the data in Table 1, the measured values ​​of organic base content in the absorption reaction solutions prepared in Examples 1-15 ranged from 29.8% to 50.3%, which basically matched the designed values ​​(30%-50%). Specifically, the organic base content in Example 6 was 29.8%, slightly lower than 30%, and in Example 7 it was 50.3%, slightly higher than 50%, both near the boundary, and the deviation was within an acceptable range. Except for Examples 2 (0.48%) and 3 (3.05%), which were close to the lower limit of 0.5% and the upper limit of 3.0% respectively, the surfactant content in the other examples ranged from 1.47% to 1.54%, with small fluctuations, indicating good repeatability of the preparation process. The moisture content was adjusted accordingly with changes in the organic base and surfactant content, and the total was close to 100%, indicating accurate formulation measurement.

[0153] The results of cloud point temperature measurements showed that the cloud points of the examples using OP-10 as the surfactant (Examples 1-8, 10-15) ranged from 68.2℃ to 76.8℃, consistent with the cloud point range of OP-10 (68-78℃). The cloud point of Example 9, using TX-10, was 64.3℃, also falling within its cloud point range (60-67℃). Regarding HLB values, the examples corresponding to OP-10 had values ​​between 13.4 and 13.9, while Example 9 corresponding to TX-10 had a value of 12.4, both conforming to their respective theoretical HLB value ranges. It is noteworthy that the cloud point of Example 2 was slightly lower (68.2℃) when the surfactant content was lower, while the cloud point of Example 3 was slightly higher (76.8℃) when the content was higher. This may be related to the effect of surfactant concentration on micelle formation temperature, but it is still within an acceptable range.

[0154] Overall, the absorption reaction solutions prepared in each embodiment meet the design requirements in terms of key performance indicators, indicating that the absorption reaction solution formulation of the present invention has good operability and stability.

[0155] Test Example 2:

[0156] The experimental steps are as follows:

[0157] (1) Take 500 mL of each of the absorption reaction solutions prepared in Example 1, Example 8, Example 9 and Comparative Example 2, and put them into a 1000 mL four-necked flask equipped with a stirrer and a gas distribution tube. Heat the flask in an oil bath and keep it at a constant temperature of 140±1℃.

[0158] (2) The reaction performance was evaluated using simulated feed gas. The simulated gas composition was: COS 1.2%, CS2 0.5%, N2 98.3%, with a total flow rate of 500 mL / min. SO2 gas was introduced at a molar ratio of 1:2 with the H2S generated after complete hydrolysis of COS and CS2 in the feed gas, i.e., the SO2 flow rate was adjusted according to real-time calculations. The gas entered the reaction liquid through the bottom gas distribution pipe after being controlled by a mass flow meter, and the stirring speed was 300 rpm.

[0159] (3) After the reaction started, samples were taken from the reactor outlet every 15 minutes using a gas bag. The concentrations of COS, CS2, H2S, SO2, and reaction products in the tail gas were analyzed using gas chromatography. Chromatographic conditions: HP-INNOWAX capillary column (30m × 0.32mm × 0.25μm), flame photometric detector (FPD), column temperature 60℃ for 2 minutes, then increased to 180℃ at 15℃ / min and held for 5 minutes. Injector temperature 200℃, detector temperature 220℃. Quantification was performed using the external standard method.

[0160] (4) The COS conversion rate is calculated as (inlet COS concentration - outlet COS concentration) / inlet COS concentration × 100%, and the CS2 conversion rate is calculated in the same way. The average value of three consecutive measurements after the reaction stabilizes (60 min later) is taken as the final conversion rate.

[0161] (5) The H2S generation rate was calculated by measuring the outlet H2S concentration and gas flow rate, in mmol / (L·min). The liquid-phase Claus reaction rate was calculated by measuring the SO2 consumption and H2S consumption per unit time, based on the stoichiometric relationship, and also expressed in mmol / (L·min). Each sample was tested in parallel three times, and the average value was taken.

[0162] (6) During the test, ensure that the volume of the absorption reaction liquid remains basically unchanged, and add a small amount of water regularly to maintain the stability of the liquid level.

[0163] The experimental results are shown in Table 2.

[0164] Table 2. Catalytic performance test results of different organic base systems:

[0165]

[0166] According to the data in Table 2, different organic base systems all exhibited catalytic activity for the hydrolysis of COS and CS2, as well as the subsequent liquid-phase Claus reaction, but the catalytic effects varied significantly. In Example 1, using pyridine as the organic base, the COS conversion rate reached 98.7%, the CS2 conversion rate was 96.5%, the H2S generation rate was 8.92 mmol / (L·min), and the liquid-phase Claus reaction rate was 4.46 mmol / (L·min). All four indicators were superior to the indole system (Example 8) and the 8-hydroxyquinoline system (Example 9). The indole system had a COS conversion rate of 97.2% and a CS2 conversion rate of 94.8%; the 8-hydroxyquinoline system had a COS conversion rate of 96.5% and a CS2 conversion rate of 93.6%. Compared to the other two systems, pyridine showed slightly higher catalytic activity, but the difference was not significant, and all three fell within the range of highly efficient catalysis.

[0167] In contrast, Comparative Example 2, using triethylamine as the organic base, showed a COS conversion of only 68.3% and a CS2 conversion of only 52.1%. The H2S formation rate and the liquid-phase Claus reaction rate were also significantly reduced, at 5.47 mmol / (L·min) and 2.73 mmol / (L·min), respectively. This indicates that although triethylamine is also an organic base, its catalytic ability for the hydrolysis of COS and CS2 is far inferior to that of pyridine, indole, and 8-hydroxyquinoline. The reason for this is that triethylamine is a tertiary amine, and the lone pair of electrons on its nitrogen atom has significant steric hindrance, making it difficult to form effective nucleophilic attack intermediates with the carbon atoms of COS and CS2. In contrast, pyridine, indole, and 8-hydroxyquinoline are all aromatic heterocyclic compounds, with the lone pair of electrons on their nitrogen atoms in a conjugated system, resulting in higher electron cloud density and stronger nucleophilicity, which is beneficial for catalyzing hydrolysis reactions.

[0168] Furthermore, in the three organic base systems, the ratio of H2S formation rate to the liquid-phase Claus reaction rate was approximately 2:1, which is consistent with the stoichiometric ratio of the Claus reaction. This indicates that the H2S generated in the liquid phase can react with SO2 in a timely manner, without H2S accumulation or reaction lag. Although this ratio is also close to 2:1 in the triethylamine system, the absolute rate is lower, indicating that its insufficient catalytic ability limits the overall reaction rate.

[0169] from Figure 2 It can be visually observed that the conversion rates and reaction rates of Examples 1, 8, and 9 are all at relatively high levels and are close to each other, while Comparative Example 2 is significantly lower. The contrast between the gray and white bars in the figure also shows that the COS conversion rate is generally slightly higher than that of CS2. This may be related to the fact that the hydrolysis of CS2 requires two steps (CS2→COS→H2S), which is kinetically slightly slower than the direct hydrolysis of COS.

[0170] In summary, the organic base selected in this invention exhibits highly efficient catalytic activity for the hydrolysis of COS and CS2 and subsequent Claus reactions, enabling the synergistic removal of multiple sulfides. The pyridine system performed best, followed by indole and 8-hydroxyquinoline, but all remained within acceptable ranges. The results of Comparative Example 2 further confirm the specificity of the selected organic base, ruling out the feasibility of conventional tertiary amine organic bases.

[0171] Test Example 3:

[0172] The experimental steps are as follows:

[0173] (1) Take 1000 mL of each of the absorption reaction solutions prepared in Example 1, Example 2, Example 3, Example 10, Comparative Example 1 and Comparative Example 4, and put them into a 2000 mL four-necked flask equipped with a stirrer and a sampling port. Heat the flask in an oil bath and keep it at a constant temperature of 140±1℃. Stir at 200 rpm.

[0174] (2) A simulated feed gas was used for the sulfur formation reaction. The simulated gas composition was: COS 1.2%, CS2 0.5%, N2 98.3%, with a total flow rate of 1000 mL / min. SO2 gas was introduced at a molar ratio of 1:2 to the H2S generated after complete hydrolysis of COS and CS2 in the feed gas. The gas was controlled by a mass flow meter and entered the reaction liquid through the bottom gas distribution pipe.

[0175] (3) After the reaction begins, every 10 minutes, use a syringe to draw about 5 mL of sample from the reaction solution and quickly inject it into a sample tube that has been preheated to 140°C. Immediately use a laser particle size analyzer to determine the particle size distribution of the liquid sulfur droplets. Take three samples each time and take the average value.

[0176] (4) Record the particle size distribution at the initial time (10 min after the reaction starts, until the reaction stabilizes), and use D50 (median volume diameter) as the characteristic particle size. Then continue to measure the particle size distribution at 20 min, 30 min, 40 min, 50 min and 60 min.

[0177] (5) The aggregation rate of sulfur droplets was calculated as (D50 at 60 min - initial D50) / 50 min, in μm / min. The particle size distribution at 60 min was recorded at the same time to observe the particle size growth.

[0178] (6) To ensure measurement accuracy, the measuring cell was cleaned with blank solvent before each sampling, and the instrument was calibrated with a standard latex ball. Each sample was tested in parallel three times, and the average value was taken.

[0179] The experimental results are shown in Table 3.

[0180] Table 3. Dispersion performance test results of different surfactant systems:

[0181]

[0182] According to the data in Table 3, different surfactant systems showed significant differences in the dispersion performance of liquid sulfur droplets. Example 1 (OP-10 1.5%) had an initial D50 of 8.2 μm, which increased to 10.9 μm after 60 min, with a coalescence rate of 0.054 μm / min. The slow particle size growth demonstrated good dispersion stability. Example 2 (OP-10 0.5%) had an initial D50 of 12.7 μm, which reached 23.6 μm after 60 min, with a coalescence rate of 0.218 μm / min. The rapid particle size growth indicated that insufficient surfactant concentration resulted in an insufficiently dense steric hindrance layer, making it difficult to effectively suppress droplet collision and coalescence. Example 3 (OP-10 3.0%) had an initial D50 of only 6.5 μm, which decreased to 8.2 μm after 60 min, with a coalescence rate of 0.034 μm / min. It exhibited the best dispersion effect, but the surfactant dosage was high, requiring a trade-off between economic factors. Example 10 (TX-10 1.5%) had an initial D50 of 9.4 μm, which increased to 14.1 μm after 60 min, with a coalescence rate of 0.094 μm / min. The dispersion effect was better than that of Example 2 but worse than that of Example 1. This is related to the lower HLB value of TX-10, which has a relatively shorter hydrophilic chain and a slightly weaker steric hindrance effect than OP-10.

[0183] Comparative Example 1 (no surfactant) had an initial D50 of 35.2 μm, which increased rapidly over time, reaching 118.4 μm after 60 min, with a coalescence rate as high as 1.664 μm / min. This indicates that without surfactant, liquid sulfur readily coalesces into large droplets, some of which may even settle and adhere to the reactor wall, increasing the risk of clogging. Comparative Example 4 (anionic surfactant sodium dodecylbenzenesulfonate) had an initial D50 of 22.8 μm, which increased to 65.3 μm after 60 min, with a coalescence rate of 0.850 μm / min. Although slightly better than the surfactant-free system, it was still far inferior to the nonionic surfactant system. This is because anionic surfactants may hydrolyze at high temperatures, and their hydrophilic heads are negatively charged, resulting in weakened electrostatic repulsion at 140°C, which fails to effectively maintain droplet stability.

[0184] from Figure 3 (a) The particle size change curve over time shows that the growth trend of Examples 1, 3 and 10 is gentle and the slope of the curve is small, while the curves of Comparative Examples 1 and 4 are steep, especially Comparative Example 1, which shows almost exponential growth in the later stage. Figure 3(b) The comparison of initial and 60-minute particle size is even more intuitive: the particle size of Comparative Example 1 after 60 minutes is as high as 3.36 times that of the initial size, Comparative Example 4 is 2.86 times, while Example 1 is only 1.33 times and Example 3 is 1.26 times. This fully demonstrates that the nonionic surfactant selected in this invention can effectively disperse liquid sulfur droplets and inhibit aggregation in the high-temperature reaction section, and OP-10 is more effective than TX-10. Within a suitable concentration range (0.5%-3.0%), the higher the concentration, the better the dispersion effect, but 1.5% can already meet industrial needs, and the benefits decrease with excessively high concentrations.

[0185] These results verify the dispersion mechanism of nonionic surfactants in the reaction stage, where they adsorb onto the sulfur droplet surface via hydrophobic groups and extend into the aqueous phase to form a steric hindrance layer. The absence of surfactants or the use of unsuitable surfactants can lead to rapid sulfur droplet aggregation, potentially causing equipment blockage and system shutdown. Therefore, the limitations imposed by this invention on the type and concentration of surfactants are of substantial significance.

[0186] Test Example 4:

[0187] The experimental steps are as follows:

[0188] (1) Collect 1000 mL of sulfur-containing solid particle slurry obtained from Examples 1-15 and Comparative Examples 1-10 after they have been stabilized under their respective process conditions, and place them in sample bottles for later use.

[0189] (2) Take 100 mL of each slurry sample and determine the particle size distribution of sulfur particles using a laser particle size analyzer. Before measurement, dilute the sample with deionized water to an appropriate concentration and ultrasonically disperse for 30 seconds to eliminate agglomeration. Record the D10, D50, and D90 values ​​(particle sizes corresponding to 10%, 50%, and 90% of the cumulative volume distribution). Each sample is measured in parallel three times, and the average value is taken.

[0190] (3) Take 200 mL of each slurry sample and pour it into a Buchner funnel (90 mm in diameter, lined with qualitative filter paper). Connect the vacuum pump and perform vacuum filtration at a vacuum of -0.08 MPa. Record the time required from the start of filtration until there is no obvious free liquid on the surface of the filter cake, which is the filtration time. Wash the filter cake twice with a small amount of deionized water and continue filtration until there is no obvious liquid dripping. Weigh the filter cake and dry it in a 105℃ oven until constant weight. Calculate the moisture content of the filter cake (weight loss after drying / wet filter cake mass × 100%). Each sample is measured in parallel 3 times and the average value is taken.

[0191] (4) Each slurry sample was subjected to a screw filter press simulation test under the filter press conditions (feed pressure 0.5 MPa) described in Examples 1-15 and Comparative Examples 1-10, and the operating torque of the filter press was recorded (expressed as a percentage of the rated torque). Each sample was measured in parallel 3 times and the average value was taken.

[0192] The experimental results are shown in Table 4.

[0193] Table 4. Test results of sulfur particle filtration performance in different embodiments and comparative examples:

[0194]

[0195] According to the data in Table 4, the sulfur particles obtained from different examples and comparative examples showed significant differences in particle size distribution, filter cake moisture content, filtration time, and filter press operating torque. In Examples 1-15, D50 was distributed between 34.2 μm and 51.3 μm, D10 between 8.5 μm and 15.2 μm, and D90 between 76.2 μm and 115.6 μm, indicating relatively uniform particle size and narrow distribution. The filter cake moisture content ranged from 9.7% to 18.5%, the filtration time ranged from 2.1 min to 5.8 min, and the operating torque ranged from 58% to 89% of the rated torque, indicating good overall filtration performance. Among them, Example 3 (OP-10 3.0%) had the largest particle size (D50=51.3μm), the lowest moisture content (9.7%), the shortest filtration time (2.1min), and the lowest torque (58%), exhibiting the best filtration performance, but with a relatively high surfactant dosage. Example 2 (OP-10 0.5%) had smaller particles (D50=28.6μm), a higher moisture content (18.5%), a longer filtration time (5.8min), and a higher torque (89%), indicating that insufficient surfactant concentration weakened the particle morphology control effect. The particle size of Examples 9 (8-hydroxyquinoline + TX-10) and 10 (TX-10) was also relatively small, and their filtration performance was slightly worse than that of the OP-10 system. This is related to the influence of surfactant type on the crystallization process. TX-10 has a lower HLB value, resulting in a slightly weaker adsorption capacity on the sulfur surface, leading to a slightly inferior crystal growth control effect.

[0196] Comparative Example 1 (no surfactant): The particles were finely broken (D50=15.3μm), with a high moisture content of 28.7%. The filtration time was 15.2 min, and the torque was 135%, making filtration extremely difficult. This indicates that without surfactant, sulfur crystals are disordered, easily forming fine particles with high moisture content. Comparative Example 5 (reaction temperature 110℃): The particles were the finest (D50=12.8μm), with the highest moisture content (32.4%). The filtration time was 18.7 min, and the torque was 152%. This indicates that when the reaction temperature is below the melting point of sulfur, sulfur precipitates in a solid state and cannot form uniform liquid droplets in the reaction section, leading to uncontrolled subsequent crystallization. Comparative Example 6 (gravity sedimentation): Although the particle distribution was similar to the example (D50=41.8μm), sedimentation separation resulted in a filter cake moisture content as high as 24.6%. The filtration time was 12.5 min, and the torque was 98%, which was far worse than filtration separation, verifying the necessity of mechanical separation. Comparative Example 7 (direct pressure filtration without cooling crystallization) exhibited extremely large particles (D50=86.4μm), very low water content (8.2%), short filtration time (1.8min), and low torque (45%), seemingly demonstrating excellent filtration performance. However, the large particles were formed by directly pressing liquid sulfur at high temperatures and then cooling it. During this process, liquid sulfur easily adheres to the equipment, posing a very high risk for long-term operation. Furthermore, the filter cake may contain uncondensed liquid sulfur, affecting product purity. Comparative Example 8 (without regeneration step) showed solvent deterioration after a certain period of operation, resulting in a worse particle distribution (D50=32.6μm), increased water content (19.8%), and a torque increase to 96%, indicating that impurity accumulation affected the regulatory function of the surfactant.

[0197] from Figure 4 (a) It can be clearly seen that the particle size distribution of Comparative Example 5 and Comparative Example 1 is significantly shifted to the left (fine particles), Comparative Example 7 is significantly shifted to the right (coarse particles), while the particle size distribution of the Examples is concentrated in the moderate range. Figure 4 (b) shows that the moisture content, filtration time, and torque of the examples are all within a reasonable range, while the comparative examples mostly deviate significantly, especially Comparative Example 5, where all indicators are extremely poor. These data confirm that in this invention, the surfactant regulates sulfur crystal growth and inhibits fine particle formation through adsorption, while a suitable cooling temperature ensures that the particles grow moderately without sticking together, thus obtaining a particle morphology that is easy to filter. Compared with gravity sedimentation, screw filtration significantly reduces separation time and filter cake moisture content, while heating clarification and regeneration ensure the continuous effectiveness of the surfactant. Therefore, the combined effect of the process conditions of this invention optimizes the filtration performance of sulfur particles and ensures long-term stable operation of the system.

[0198] Test Example 5:

[0199] The experimental steps are as follows:

[0200] (1) Take the purified gas from the top outlet of the absorption reaction tower of Examples 1-15 and Comparative Examples 1-10 after they have been running continuously and stably under their respective process conditions (running time ≥72h) and take samples with a gas bag. Take samples of each sample once every 15 minutes within 1 hour, for a total of 4 times, mix them evenly and use them as the test samples.

[0201] (2) The COS, CS2, H2S, and total sulfur content in the purified gas were analyzed using gas chromatography. Chromatographic conditions: HP-INNOWAX capillary column (30m × 0.32mm × 0.25μm), flame photometric detector (FPD), equipped with a sulfur filter. Column temperature was programmed: initial temperature 60℃, held for 2 min, then increased to 180℃ at 15℃ / min and held for 5 min. Injector temperature 200℃, detector temperature 220℃. Carrier gas was high-purity nitrogen, flow rate 1.5mL / min, split ratio 10:1. Injection method: automatic gas injection valve, quantitative loop volume 1mL.

[0202] (3) Quantification was performed using the external standard method; a series of standard gases of different concentrations of COS, CS2, and H2S (with nitrogen as the base gas) were prepared, and the peak areas were measured under the same chromatographic conditions to plot a standard curve. For sample determination, qualitative analysis was performed based on retention time, and quantitative analysis was performed based on peak area. Each sample was measured in triplicate, and the average value was taken.

[0203] (4) The total sulfur removal rate is calculated as (total sulfur content at the inlet - total sulfur content at the outlet) / total sulfur content at the inlet × 100%. The total sulfur content at the inlet is calculated from the composition of the feed gas (COS and CS2 content), wherein the COS content is 1.2% and the CS2 content is 0.5% (Examples 1-13, Comparative Examples 1-10) or adjusted accordingly based on the feed gas composition of Examples 14-15. The feed gas composition in the comparative examples is the same as that in Example 1, unless otherwise stated.

[0204] The experimental results are shown in Table 5.

[0205] Table 5. Sulfur content and removal rate of purified gas in different embodiments and comparative examples:

[0206]

[0207] Note: ① In Example 14, the total sulfur content (volume fraction) of the raw gas was 0.7%, and the total sulfur concentration at the inlet was 7000 mg / m³. 3 Conversion (standard conditions); ② Example 15: Total sulfur content of feed gas is 3.5%, and the total sulfur concentration at the inlet is 35000 mg / m³. 3 Conversion; the total sulfur content of the feed gas in the other embodiments and comparative examples is 1.7%, and the total sulfur concentration at the inlet is 20000 mg / m³. 3 Conversion. The above conversion values ​​are based on the simplified assumption that 1% by volume of sulfur-containing compounds corresponds to 10,000 mg / m³.3 This allows for a more intuitive comparison; the theoretical value calculated using the standard gas law (22.4 L / mol) would be higher, but the ratio of inlet to outlet concentration remains unchanged in the removal rate calculation, so the removal rate data is unaffected.

[0208] According to the data in Table 5, the total sulfur content of the purified gas in Examples 1-15 was all around 4.1 mg / m³. 3 ~15.5mg / m 3 The total sulfur removal rate was greater than 99.2% in all cases, with Example 3 (surfactant concentration 3.0%) showing the lowest total sulfur content (4.1 mg / m³). 3 Example 1 (0.5% surfactant) had the highest removal rate (99.79%); Example 2 (0.5% surfactant) had the highest total sulfur content (15.5 mg / m³). 3 The removal rate was 99.23%, still meeting the purification requirements. Example 9 (8-hydroxyquinoline + TX-10) had a total sulfur content of 13.8 mg / m³. 3 The removal rate was 99.31%, slightly lower than that of the pyridine + OP-10 system, which is consistent with the previous catalytic performance and dispersion performance test results. Examples 14 and 15 treated low-sulfur and high-sulfur feed gases, respectively, with removal rates of 99.72% and 99.67%, indicating that the method of the present invention has good adaptability to feed gases with different sulfur loads.

[0209] In the comparative examples, Comparative Example 1 (without surfactant) had a total sulfur content of 86.3 mg / m³. 3 The removal rate was 95.69%, far lower than in the examples, indicating that surfactants are crucial for ensuring desulfurization efficiency. Comparative Examples 2 (triethylamine) and 4 (anionic surfactant) had removal rates of 96.93% and 96.51%, respectively, both unsatisfactory. Comparative Example 5 (reaction temperature 110°C) had a removal rate of only 92.88%, with a total sulfur content as high as 142.5 mg / m³, proving that the reaction temperature being below the sulfur melting point led to a severe decrease in reaction efficiency. Comparative Example 6 (gravity sedimentation) had a removal rate of 99.54%, similar to the examples, but previous tests of Example 4 showed poor filtration performance and unreliable long-term operation. Comparative Example 7 (no cooling crystallization) had a removal rate of 99.75%, even higher than most examples, but as mentioned earlier, its operation risk is high, limiting its practical industrial application value. Comparative Example 8 (no regeneration step) had a removal rate of 98.88%, with a total sulfur content of 22.4 mg / m³. 3 The results were significantly worse than those of the examples, indicating that impurity accumulation affected the reaction efficiency. The removal rates of Comparative Example 9 (low organic base concentration) and Comparative Example 10 (low regeneration temperature) were 97.89% and 98.55%, respectively, both lower than those of the examples of this invention.

[0210] These data demonstrate the high efficiency of the method of this invention in the synergistic removal of multiple sulfides. The fundamental reason lies in the dual catalytic effect of the organic base, which ensures the complete hydrolysis of COS and CS2 and the rapid reaction of H2S and SO2; the dispersing effect of the surfactant prevents the reduction of the reaction interface caused by sulfur polymerization; the appropriate temperature gradient ensures the coordinated progress of reaction and separation; and heating, clarification, and regeneration maintain the stability of solvent activity. The synergistic effect of each step ultimately achieves an excellent total sulfur removal rate of >99.5%.

[0211] Test Example 6:

[0212] The experimental steps are as follows:

[0213] (1) Collect approximately 500g of sulfur products (after drying) obtained from Examples 1-15 and Comparative Examples 1-10 after stabilizing under their respective process conditions, place them in sealed sample bags, and label them for later use.

[0214] (2) Weigh approximately 0.2 g (accurate to 0.0001 g) of the ground sulfur sample and place it in a porcelain boat of known mass. Burn the sample in an oxygen stream until all the sulfur is converted into sulfur dioxide. Absorb the sulfur dioxide with hydrogen peroxide absorbent and oxidize it into sulfuric acid. Titrate the sulfur dioxide with a standard sodium hydroxide solution and calculate the sulfur content based on the amount consumed. Perform three parallel determinations for each sample and take the average value.

[0215] (3) Weigh approximately 25g of sulfur sample (accurate to 0.01g), place it in a pre-weighed porcelain crucible, and slowly heat it on an electric furnace to allow the sulfur to completely volatilize. Then, ignite it in a high-temperature furnace at 800±25℃ until constant weight, and calculate the ash content from the mass of the residue. Each sample is measured in parallel three times, and the average value is taken.

[0216] (4) Weigh about 25g of sulfur sample (accurate to 0.01g), add 50mL of carbon dioxide-free water, heat to boiling for 5min, cool and filter, add phenolphthalein indicator to the filtrate, and titrate with sodium hydroxide standard solution until a light pink color is obtained, and calculate the acidity (as H2SO4). Each sample is measured in parallel 3 times and the average value is taken.

[0217] (5) Weigh approximately 10g of sulfur sample (accurate to 0.001g), place it in a pre-weighed weighing bottle, dry it in an oven at 80±2℃ for 3h, remove it, cool it in a desiccator, weigh it, and calculate the moisture content. Each sample is measured in parallel 3 times, and the average value is taken.

[0218] (6) Weigh approximately 10g of sulfur sample (accurate to 0.001g) and place it in a porcelain boat of known mass. Heat the sample at 250±10℃ to melt the sulfur and allow it to slowly volatilize. Burn the residue in a high-temperature furnace at 800±25℃ until constant weight. Calculate the organic content based on the mass lost during burning of the residual organic matter. Perform three parallel determinations for each sample and take the average value.

[0219] The experimental results are shown in Table 6.

[0220] Table 6. Quality test results of sulfur products from different embodiments and comparative examples:

[0221]

[0222] According to the data in Table 6, the purity of the sulfur products obtained in Examples 1 to 15 is all above 99.38%, with Example 3 (surfactant 3.0%) having the highest purity (99.81%) and Example 9 (8-hydroxyquinoline + TX-10) having the lowest purity (99.38%), but still meeting the requirements for first-grade industrial sulfur (≥99.5%). The ash content, acidity, moisture, and organic matter content are all at low levels: ash 0.021%–0.064%, acidity 0.0028%–0.0094%, moisture 0.08%–0.24%, and organic matter 0.031%–0.085%, all exceeding the first-grade standards of GB / T 2449-2006 (ash ≤0.10%, acidity ≤0.01%, moisture ≤0.20%, organic matter ≤0.30%). Some indicators in Examples 2 and 9 (such as moisture 0.21% and organic matter 0.076%) were slightly higher, but still qualified.

[0223] In the comparative examples, Comparative Example 1 (no surfactant) had a sulfur purity of only 98.25%, with significantly higher ash, acidity, and organic matter content, at 0.187% and 0.231% respectively, far exceeding the standard. Comparative Examples 2 (triethylamine), 4 (anionic surfactant), and 5 (low-temperature reaction) all had sulfur purities below 99%, with high impurity content. Comparative Example 5 had an ash content as high as 0.235% and organic matter 0.296%, close to the upper limit of the standard. Comparative Examples 6 (gravity sedimentation) and 7 (no cooling crystallization) had sulfur purities of 99.12% and 99.35% respectively, slightly lower than the examples but still within the standard, but based on previous tests, their operational stability was poor. Comparative Example 8 (no regeneration step) had a sulfur purity of 98.92%, with a significant increase in impurities. Comparative Examples 9 (low organic alkali concentration) and 10 (low regeneration temperature) had sulfur purities of 98.68% and 99.08% respectively, with also high impurity content.

[0224] These data demonstrate that the method of this invention can produce high-purity, low-impurity sulfur products. This is because the dispersing effect of the surfactant prevents sulfur particles from encapsulating organic alkalis and impurities; the aromatic washing after pressure filtration effectively removes residual organic alkalis and surfactants from the filter cake; and the heating clarification and regeneration process removes impurities from the circulating liquid through cloud point separation, preventing them from entering the sulfur product. Therefore, this invention achieves efficient desulfurization while ensuring the quality of the sulfur product.

[0225] Test Example 7:

[0226] The experimental steps are as follows:

[0227] (1) Take the systems of Example 1 and Comparative Example 8, which are running continuously under their respective process conditions, and the control experiment (hereinafter referred to as the control experiment) which follows the same formula and process as Example 1 but eliminates the correlation control between the filter pressure torque and the temperature of the heating clarification tank and keeps the temperature of the heating clarification tank constant at 95°C. Set up sampling ports on the pipeline from the circulating regeneration device back to the absorption reaction tower. When the system running time reaches 500 hours, 1000 hours and 2000 hours, take 100 mL samples, place them in sample bottles, and label them for later use.

[0228] (2) The content of organic bases was determined by gas chromatography under the same chromatographic conditions as in Test Example 1. 1 μL of the sample was directly injected after filtration through a 0.45 μm filter membrane. Quantification was performed using the external standard method, and the pyridine content was calculated. Each sample was measured in triplicate, and the average value was taken.

[0229] (3) The surfactant content was determined by ultraviolet spectrophotometry under the same conditions as in Test Example 1. The sample was diluted with distilled water to an appropriate concentration, and the absorbance was measured at 275 nm. The OP-10 content was calculated based on the standard curve. Each sample was measured in triplicate, and the average value was taken.

[0230] (4) Total organic carbon (TOC) was determined using a TOC analyzer. After filtering the sample through a 0.45 μm filter membrane, the TOC value was measured using the total organic carbon mode and recorded. Each sample was measured in triplicate, and the average value was taken.

[0231] (5) Qualitative analysis of impurity types and contents was performed using gas chromatography-mass spectrometry (GC-MS). Chromatographic conditions: HP-INNOWAX capillary column (30m×0.32mm×0.25μm), injection port temperature 250℃, column temperature programmed: initial temperature 50℃ held for 2 min, then increased to 250℃ at 10℃ / min and held for 10 min. Mass spectrometry conditions: EI ion source, 70 eV, scan range m / z 35-500. 1 μL of sample was extracted with dichloromethane and injected. Impurity types were identified by comparison with the NIST spectral library. The total impurity content (expressed as a percentage of peak area relative to pyridine) was estimated semi-quantitatively using peak area normalization.

[0232] (6) Solvent replenishment is calculated as (initial organic base mass - current organic base mass) / operating time, in kg / month (converted to 30 days). Solvent lifespan is estimated based on the time required for the organic base content to drop to 80% of its initial value.

[0233] The experimental results are shown in Table 7.

[0234] Table 7. Solution stability test results for different systems:

[0235]

[0236] According to the data in Table 7, Example 1, Comparative Example 8, and the control experiment exhibited significantly different solution stability during long-term cyclic operation. After 2000 hours of operation, Example 1 showed a decrease in organic base content from the initial 40.2% to 38.5%, a decrease of 4.2%; surfactant content decreased from 1.52% to 1.36%, a decrease of 10.5%; TOC increased from 1820 mg / L to 2450 mg / L; the number of impurity types increased from 0 to 9; and the total impurity content (peak area %) reached 5.6%. Solvent replenishment increased slightly with operating time, from an average of 0.58 kg / month over 500 hours to an average of 0.67 kg / month over 2000 hours, with an estimated solvent lifespan of approximately 8600 hours. These data indicate that the solution in Example 1 maintained good stability during long-term operation, with slow loss of organic base and surfactant, and effective control of impurity accumulation.

[0237] In Comparative Example 8 (without a regeneration step), after 2000 hours of operation, the organic base content plummeted from 40.1% to 29.5%, a decrease of 26.5%; the surfactant content decreased from 1.51% to 0.52%, a decrease of 65.6%; TOC surged from 1840 mg / L to 8920 mg / L, the number of impurity types increased to 25, and the total impurity content reached as high as 28.5%. Solvent replenishment increased from 1.52 kg / month for 500 hours to 2.23 kg / month for 2000 hours, approximately 3-4 times that of Example 1, and the solvent lifespan was only 1350 hours. These data clearly demonstrate that without a regeneration step, impurities continuously accumulate in the system, not only diluting and consuming the effective components but also leading to accelerated degradation of organic bases and surfactants, resulting in rapid solution deterioration.

[0238] The control experiment (without associated controls) showed performance between the two. After 2000 hours of operation, the organic base content decreased to 35.1%, a decrease of 12.9%; the surfactant content decreased to 1.04%, a decrease of 32.0%; TOC increased to 4520 mg / L, with 19 types of impurities and a total impurity content of 17.3%. The solvent replenishment rate increased from 0.86 kg / month for 500 hours to 1.38 kg / month for 2000 hours, approximately twice that of Example 1, and the solvent lifespan was 3800 hours. This indicates that although the control experiment used heated clarification regeneration, the lack of associated control between the pressure filtration torque and regeneration temperature prevented dynamic adjustment of the regeneration intensity based on the system status, resulting in incomplete impurity removal or excessive organic base loss, and stability inferior to Example 1.

[0239] from Figure 5 The divergence trend of the three curves can be seen intuitively. Figure 5In (a), the organic base and surfactant in Example 1 decreased gradually, while the decrease in Comparative Example 8 was steep, and the control experiment was in the middle. Figure 5 In (b), the TOC and impurity content of Example 1 increased slowly, while that of Comparative Example 8 showed an exponential growth trend.

[0240] These data validate the two key innovative mechanisms of this invention. First, the heating-based clarification and regeneration utilizes the cloud point characteristics of surfactants to enrich and remove impurities, effectively preventing their accumulation in the system. Second, the correlation control between the pressure filtration torque and the regeneration temperature enables dynamic optimization of the regeneration intensity based on the system state, ensuring impurity removal while minimizing organic alkali loss. The synergistic effect of these two mechanisms allows the solution to maintain a stable composition and activity during long-term operation, thus guaranteeing the long-term stable operation of the system.

[0241] Test Example 8:

[0242] The experimental steps are as follows:

[0243] (1) Samples were taken from the systems of Examples 1, 14, and 15, which were running continuously under their respective process conditions, and tested after the systems had stabilized (running time ≥ 72h). Each example maintained its specific feed gas composition: Example 1 used the reference feed gas (COS 1.2%, CS2 0.5%, total sulfur content approximately 20000mg / m³). 3 Example 14 uses low-sulfur feed gas (COS 0.5%, CS2 0.2%, total sulfur content approximately 7000 mg / m³). 3 Example 15 uses high-sulfur feed gas (COS 2.5%, CS2 1.0%, total sulfur content approximately 35000 mg / m³). 3 ).

[0244] (2) Determination of total sulfur removal rate: A gas bag was used to collect a sample from the purified gas sampling port at the top outlet of the absorption reaction tower. The COS, CS2, H2S and total sulfur content in the purified gas were analyzed using a gas chromatograph under the same chromatographic conditions as in Test Example 5. The total sulfur removal rate was calculated based on the inlet total sulfur concentration and the outlet total sulfur concentration. Each example was measured continuously for 3 days, with 3 samples taken each day, and the average value was taken.

[0245] (3) Determination of sulfur product purity: Collect approximately 500g of sulfur product (after drying) obtained from each example and determine the sulfur purity according to the method in Test Example 6. Each sample was measured in parallel 3 times and the average value was taken.

[0246] (4) Determination of filtration performance: Take 500 mL of sulfur-containing solid particle slurry obtained after each example has been stabilized, and determine the D50 particle size of the sulfur particles and the moisture content of the filter cake according to the method described in Test Example 4. Each sample is measured in parallel 3 times, and the average value is taken.

[0247] (5) Determination of solvent stability: After 500 hours of continuous operation in each embodiment, samples were taken from the pipeline returning to the absorption reaction tower from the recycling regeneration unit, and the organic base content, surfactant content, and TOC were determined according to the method described in Test Example 7. The percentage deviation from the initial value was calculated. Each sample was measured in triplicate, and the average value was taken.

[0248] The experimental results are shown in Table 8.

[0249] Table 8. Results of adaptability tests under different feed gas conditions:

[0250]

[0251] According to the data in Table 8, Examples 1, 14 and 15 were operated under baseline, low-sulfur and high-sulfur feed gas conditions, respectively, and all performance indicators were good, indicating that the method of the present invention has strong adaptability to feed gas with different sulfur loads.

[0252] In terms of desulfurization efficiency, the total sulfur removal rate of all three examples was above 99.6%. The low-sulfur feed gas (Example 14) had the highest removal rate (99.72%), while the high-sulfur feed gas (Example 15) had a slightly lower removal rate (99.67%), but they remained at the same level. The total sulfur content at the outlet increased with the increase of the inlet sulfur load, from 5.5 mg / m³. 3 (Low sulfur) up to 7.6 mg / m³ 3 (Baseline) to 11.6 mg / m 3 (High sulfur content), but all levels are far below environmental emission limits. This indicates that the dual catalytic mechanism of the organic base in this invention can effectively function under different sulfur loads: the hydrolysis reaction rates of COS and CS2 are positively correlated with the reactant concentrations. Under high sulfur loads, the amount of H2S generated per unit time is greater, which places higher demands on the processing capacity of the liquid-phase Claus reaction system. However, the results show that the reaction can still proceed completely without breakthrough or efficiency reduction.

[0253] Regarding the purity of the sulfur products, all three examples showed a purity between 99.66% and 99.70%, with no significant difference, and all met the requirements for first-grade industrial sulfur. This indicates that the quality of the sulfur products does not fluctuate due to changes in the sulfur load of the raw material gas, and that the aromatic washing after pressure filtration effectively removes residual organic alkalis and surfactants from the filter cake, ensuring stable product purity.

[0254] In terms of filtration performance, the D50 of sulfur particles increased slightly with increasing sulfur load, from 41.5 μm for low sulfur to 42.7 μm for the baseline and then to 43.2 μm for high sulfur. The filter cake moisture content showed the opposite trend: 13.3% for low sulfur, 12.8% for the baseline, and 12.5% ​​for high sulfur, indicating that the larger particles formed a filter cake with higher porosity and more thorough dewatering. Overall, the filtration performance was within a good range, and the filter press operated stably.

[0255] Regarding solvent stability, after 500 hours of operation, the loss of organic bases and surfactants, as well as the increase in TOC, all intensified with increasing sulfur load. Under high-sulfur feed gas conditions, the organic base content decreased by 1.8%, surfactants decreased by 4.2%, and TOC increased by 10.8%; while under low-sulfur conditions, these figures were only 0.8%, 1.9%, and 4.2%, respectively. Under baseline conditions, the organic base content decreased by 1.0%, surfactants decreased by 2.6%, and TOC increased by 7.1%, falling between the two. This is in line with expectations: a higher sulfur load means a larger total amount of sulfides processed per unit time, resulting in more byproducts and impurities generated during the reaction, naturally exacerbating solvent contamination and loss. However, even so, the changes after 500 hours under high-sulfur conditions remained within acceptable limits, and the solvent could maintain effective operation through heating clarification and regeneration, as well as dynamic correlation control.

[0256] In summary, the method of this invention maintains stable desulfurization efficiency, product quality, and filtration performance across a wide range of sulfur content in the feed gas. While solvent stability varies, it remains within a controllable range. This is attributed to the stable system comprised of the highly efficient catalytic effect of the organic base, the dispersion and regulation function of the surfactant, and the impurity removal mechanism through heating, clarification, and regeneration. Therefore, the method of this invention has good adaptability to feed gas and can be widely applied to the treatment of industrial exhaust gases from various sources.

[0257] Although embodiments of the invention have been shown and described, it will be understood by those skilled in the art that various changes, modifications, substitutions and alterations can be made to these embodiments without departing from the principles and spirit of the invention, the scope of which is defined by the appended claims and their equivalents.

Claims

1. A sulfur recovery method based on organic base catalytic hydrolysis and spiral pressure filtration, characterized in that, Includes the following steps: S1. The raw gas containing COS and / or CS2 and SO2 gas are introduced into an absorption reaction tower containing an absorption reaction liquid and reacted at 120-150°C to hydrolyze COS and / or CS2 to generate H2S. The generated H2S reacts with SO2 in a liquid-phase Claus reaction to generate elemental sulfur. The absorption reaction liquid contains an organic base, an aromatic solvent, a nonionic surfactant and water. S2. Cool the absorbent reaction liquid containing liquid sulfur obtained after the reaction to below the freezing point of sulfur to obtain a slurry containing sulfur solid particles. S3. The slurry is fed into a screw filter press for solid-liquid separation to obtain sulfur filter cake and filtrate; S4. The filtrate is heated to above the cloud point temperature of the nonionic surfactant, allowed to stand and separate into layers, and the upper layer of regenerated organic alkaline aqueous solution and the lower layer of heavy phase rich in surfactant and impurities are obtained. S5. Return the regenerated organic alkaline aqueous solution from the upper layer to the absorption reaction tower for recycling. The nonionic surfactant is used to disperse liquid sulfur droplets in the contact reaction, to regulate the morphology of sulfur particles in the cooling step, and to separate and enrich impurities by cloud point separation in the heating and static stratification steps. The organic base is pyridine, indole, or 8-hydroxyquinoline; the aromatic solvent is tricresylbenzene or heavy aromatic hydrocarbons; the nonionic surfactant is alkylphenol polyoxyethylene ether with a cloud point of 60–78°C and an HLB value of 12–14.

2. The sulfur recovery method based on organic base catalytic hydrolysis and spiral pressure filtration according to claim 1, characterized in that, In the absorption reaction solution, the mass percentage of organic base is 30% to 50%, the mass percentage of aromatic solvent is 5% to 10%, the mass percentage of nonionic surfactant is 0.5% to 3.0%, and the balance is water.

3. The sulfur recovery method based on organic base catalytic hydrolysis and spiral pressure filtration according to claim 1, characterized in that, The reaction temperature inside the absorption reaction tower is 130–150°C; the cooling temperature in the cooling step is 50–70°C; and the heating temperature in the filtrate heating step is 90–98°C.

4. The sulfur recovery method based on organic base catalytic hydrolysis and spiral pressure filtration according to claim 1, characterized in that, The molar ratio of the SO2 gas flow rate to the H2S generated after complete hydrolysis of COS and / or CS2 in the feed gas is 1:1.8 to 1:2.

2.

5. The sulfur recovery method based on organic base catalytic hydrolysis and spiral pressure filtration according to claim 1, characterized in that, The feed pressure of the spiral filter press is 0.3 to 0.8 MPa, the spiral compression ratio is 2:1 to 4:1, and the internal temperature is maintained below the melting point of sulfur by a cooling jacket during the filtration process.

6. The sulfur recovery method based on organic base catalytic hydrolysis and spiral pressure filtration according to claim 1, characterized in that, The sulfur filter cake separated by the spiral filter press is reverse washed with an aromatic solvent at the end of the filter press, and then dried to obtain the sulfur product.

7. The sulfur recovery method based on organic base catalytic hydrolysis and spiral pressure filtration according to claim 1, characterized in that, The content of COS in the raw gas is 0.1% to 5%, and the content of CS2 is 0.05% to 3%.

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