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I.CHEM.E.SYMPOSIUM SERIES NO. 110

EMERGENCY RELIEF REQUIREMENTS FOR REACTIVE CHEMICAL STORAGE TANKS

J.C.Leung*, A.C.Buckland**, A.R.Jones**, L.D.Pesce

This paper presents a detailed case study on the emergency venting requirements for hydrogen cyanide (HCN), with a potential for application to other reactive chemicals. The concern focuses on the uncontrolled polymerization of HCN in storage facilities. Both thermal data and flow regime data during venting were obtained using the DIERS bench-scale apparatus for characterizing runaway reactions. This study points out the importance of (1) detecting the presence of a continued runaway after relief venting and (2) providing planned actions, such as a reaction quench, to mitigate the consequences.

INTRODUCTION

This work was first initiated three years ago by two U.S. companies and later expanded in scope to involve participation by about ten major chemical companies worldwide. Though this is a detailed case study addressing the venting requirements for HCN storage facilities, the experimental technique and the vent sizing methodology presented here would be generally applicable to other reactive chemical storage systems.

Background

HCN is used mainly in the manufacturing of nylon and methyl methacry-late plastics. In the U.S., the 1987 demand for HCN was over one billion pounds. This large industrial demand results in the need for some storage of HCN. HCN is stored as a liquid (NBP 25.7°C) in refrigerated tanks which are protected from excessive pressure buildup by emergency relief devices. Prior to this study, these relief devices were usually sized to protect the storage tanks from overpressure caused by fire.

HCN is subject to polymerization (Gause and Montgomery [1]) similar to many monomeric chemicals. The heat of polymerization further accelerates the reaction leading to runaway situations (Walker and Eldred [2]). Under normal storage HCN is stabilized with acid to prevent polymer formation. Accidental uncontrolled polymerizations would be caused by contamination of the HCN by any material which will either consume the stabilizer or increase the pH of the HCN to the 5 to 10 range. Uncontaminated HCN would only undergo runaway polymerization at high temperatures (> 100°C).

*Fauske & Associates, Inc., Burr Ridge, Illinois, U.S.A..

**American Cyanamid, Westwego, Louisiana, U.S.A.

E. I. DuPont de Nemours & Company, Memphis, Tennessee, U.S.A.

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I.CHEM.E. SYMPOSIUM SERIES NO. 110

EXPERIMENTAL SET-UP

The predecessor of the commercial apparatus, Vent Sizing Package (VSP was developed by Fauske & Associates, Inc. and is currently marketed by Fike Corporation, Blue Springs, Missouri, U.S.A.), was the DIERS (Design Institute for Emergency Relief Systems) bench-scale apparatus which has been described by Fauske and Leung [3], and Leung et al [4]. The low thermal mass test cell (SS-304) with a volume of 120 ml is contained in a 4 liter containment vessel as shown in Figure 1. Measurements consist of sample temperature (Tl) and pressure (PI), external guard temperature (T2), and containment pressure (P2). During the runaway or self-heating period, the guard heater assembly serves to provide an adiabatic environment for the test sample by regulating T2 close to Tl. For closed system adiabatic tests, the containment vessel serves to prevent bursting of the test cell by regulating its own pressure to follow the test cell pressure. For open system tests, pressure communication with the containment is provided by a top vent (usually 1-2 mm in diameter) in the test cell. In most cases, an initial back pressure usually determined by the desired relief set pressure is imposed on the system.

Modification

Flow regime tests are useful in discerning between foamy (bubbly) and non-foamy (churn-turbulent) character of the sample during venting. Here instead of directing the vented material into the containment vessel, a slight modification is made whereby the effluent is discharged to the outside as shown in Figure 1. The test cell in this configuration is initially sealed and when the desired relief pressure or temperature is reached, the full-flow valve in the discharge line is opened to initiate the blowdown.

Another modification to the normal set-up is aimed at better understanding of the reaction rate phenomena during the post-relief period. Here the test cell and the heater assembly are taken out of the containment vessel and placed on an electronic weight scale to monitor the mass loss. In addition, pressure drop across the top vent is measured by a highly sensitive DP cell. This external set-up as depicted in Figure 2 focuses on the vapor streaming rate, as a measure of the reaction energy release during the tempering phase, and the vapor/gas production rate during the second exotherm as discussed later.

WORST CASE SCENARIO

At the outset, it seems impossible to define the worst case contamination mixture. Here one is confronted with a wide variety of possible contaminants which sometimes are plant specific. In the first category, we have the alkaline chemicals such as caustic, ammonia, and ammonium hydroxide which are present in process streams. In the second category, we have the coolant medium such as a mixture of water and methanol. In the last category, we have the in-situ produced HCN polymer which can accumulate in dead spaces particularly if the tanks are not washed out thoroughly. All of the above-mentioned initiators have been tested in this study. As a possible substitute for anhydrous ammonia, triethylamine (TEA is a liquid at ambient conditions) was also used.

Since the contaminant level is expected to influence the polymerization rate, we have to decide the amount that could give rise to the worst case. Here one is faced with the dilemma of defining what levels constitute credible, or incredible contamination. This problem was circumvented in this study by first carrying out the test with a seemingly incredible amount of contaminant which also gave rise to a total reaction heat release rate

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similar in magnitude as the anticipated fire load. Using an 80 m (21,000 gallon) vessel and the API formula [5], the fire load is calculated to yield a temperature rise of about l°C/min. The standard contaminant based on 30% wt. caustic solution and at a level equal to 18% wt. of the HCN present (this is equivalent to a NaOH or NaCN concentration of 1.68 g-mole/kg HCN, or 1.37 g-mole/kg mix) was found to yield similar self-heat rates at relief set pressure of 0.34 bar g (5 psig). Though this massive contamination scenario may approach the incredible worst case, the emergency venting requirement is of the same order of magnitude as the fire exposure situation.

RUNAWAY CHARACTERIZATION

Closed System Behavior

The runaway behavior of contaminated HCN is best illustrated by a closed system test. Though the contaminant level was less than the standard case, the resulting self-heat rates were not much different in the region of interest. Figure 3 illustrates the temperature and pressure behavior. Note that the runaway was initiated at room temperature. Near 100°C, a very rapid excursion took place, leading to eventual rupture of the test cell. Figure 4 shews that the rates of temperature rise, commonly known as self-heat rates, are in accordance with the Arrhenius behavior. Figure 4 also illustrates that the pressure-temperature data closely follow the HCN vapor pressure curve; the slight off-set can be attributed to the water present.

Open System Behavior

It is equally important to examine the runaway behavior in an open system whereby the reaction heat can be removed via vaporization of the volatile components; the so-called tempering condition is attained with no further rise in temperature. Figure 5 illustrates the tempering conditions achieved at both 1 bar g and 0 bar g. Note that following the tempering period, the sample exhibits a final temperature excursion which far exceeds the rates inferred from the closed system data. The following points are worth noting:

1. The delay time or the tempering period is significantly reduced at the higher tempering temperature. The latter is caused by the higher imposed back pressure.

2. The temperature rise rate during the final excursion is also greater for the case at the higher tempering temperature.

3. Based on (1) and (2), it would appear that a non-reclosure type relief device such as a rupture disk is preferred. This would allow the reaction to temper at the lowest possible temperature corresponding to 0 bar g and take advantage of the longer delay time.

The remaining question is what is the relative venting requirement between the earlier exotherm and this last exotherm. Furthermore, is it feasible to design a practical relief system to cope with this later exotherm? To answer this question, additional tests aimed at measuring vapor/gas generation rate were carried out as discussed in the next section.

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More Detailed Measurements

Detailed measurements involving weight loss of the sample and pressure drop across a calibrated vent tubing were made during the tempering phase and the final temperature excursion. The arrangement as shown in Figure 2 was set up inside the hood with adequate shielding. Since we are most interested in the phenomena occurring at ambient pressure following a postulated rupture-disk relief, the particular set-up is most convenient without the containment vessel. The 1.08 mm diameter vent line was calibrated by boiling water in the test cell while noting the mass loss rate and the pressure drop. This calibration yielded a discharge coefficient of 0.85.

Figure 6 presents the data obtained with TEA contaminated HCN showing similar delay time. It is interesting to note that during the tempering period the reaction rate was actually increasing as inferred from the weight measurement. During this period the HCN vapor flow rate was insufficient to give any noticeable pressure drop in the vent line. However, simultaneous with the last temperature excursion, a significant pressure drop was detected. By finite differencing the weight data, the mass flow rate was also found to exhibit a peak at the same instance. Using the Bernoulli equation in the f o r m , we can rewrite to express the fluid density as

Taking the following values,

the density is calculated to be 1.3 kg/m which is very close to the vapor density of HCN (1.1 kg/m ) at 0 bar g. These same data will be examined in terms of venting requirement later on.

In principle, one should be able to obtain the reaction rate data based on the measured mass loss rate together with information on the latent heat of vaporization and the heat of polymerization. However, this kind of detailed kinetic modeling is beyond the scope of this paper.

ACID QUENCH STUDY

Several tests were conducted to test the feasibility of quenching the reaction once polymerization had started as an added safety measure in pre-relief and post-relief considerations. Being able to quench the reaction would potentially reduce emergency relief venting and prevent the last excursion at the end of the tempering period. As it turns out, mitigation of this last exotherm by acid quench is most essential.

The quenching studies consisted of adding acids (sulfuric acid, acetic acid, sulfur dioxide, phosphoric acid, and hypophosphorous acid) and water in amounts large enough to quench the reaction at 25-50°C and provide long term stabilization of the contaminated HCN. Intuitively the amount of acid added should exceed the minimum requirement for neutralization with the base contaminant. The acid equivalency for neutralization was determined

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experimentally with the use of a pH meter and the results shown in Table 1 provide a practical guideline for minimum acid quench requirement. Sulfuric acid is a strong acid, hence it is most efficient on a per mole basis. However, its usage does require adequate dilution with water so as to avoid the hydrolysis reaction with HCN locally. Sulfur dioxide also presents itself as a good quench agent but it might require special distribution system for its effective dispersion.

Reaction quench was indeed feasible as demonstrated by the prolonged tempering or temperature stability at around 30°C using various acid quenches. As a rule of thumb, effective quenching seems best accomplished by using twice the amount of acid needed for neutralization. Finally, it should be mentioned that dilution with equal quantity of water as the contaminated HCN was found to lessen the severity of the last excursion to tolerable range (maximum rate was 5°C/min as compared to over 100°C/min without water). Apparently the dilution and the heat capacity effects in this case outweigh the instability normally induced by water at lower concentration. Furthermore, if the contaminant is of the type that will form an immiscible layer with HCN (such as strong caustic), then in order to have an effective quench, enough water must be added to render the layers miscible.

VENTING REQUIREMENTS

Flow Regime Consideration

For emergency relief situations, the DIERS program points out the necessity of considering two-phase flow in the vent line (Fisher [6]). However, there exist varying degrees of two-phase venting depending on the in-vessel flow regime behavior. Typically these flow regimes can be classified as:

• homogeneous or uniform froth,

• bubbly,

• churn-turbulent.

The first regime is an idealized situation where no vapor-liquid disengagement is assumed. Bubbly regime is indicative of relatively little vapor disengagement. Generally, there is little distinction between the bubbly regime and the homogeneous regime; both are foamy in nature. In contrast, for the churn-turbulent regime, relatively more vapor disengagement is possible, i.e., the vapor mass fraction or quality entering the vent can differ substantially from the vessel average quality.

Discrimination between foamy and non-foamy behavior is of special interest in this study, not just from a vent sizing point of view but also from an effluent control consideration. The flow regime test was carried out by venting the mixture through a 2.4 mm vent from an initial pressure of 15 psig. For other systems, this size vent was found to leave behind 40-60% of the liquid if non-foamy behavior prevails and essentially no liquid at all if foamy regime prevails. Two-phase flow, as indicated by the presence of entrained liquid exiting from the vent line, was observed. Table 2 compares the overall flow regime result with other systems; clearly HCN exhibits non-foamy (churn-turbulent) behavior during venting under runaway conditions.

As an additional consideration, a simulation test was conducted using a vent line of 1.08 mm in diameter. This would yield equivalent superficial

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2 2 velocities for a vessel with a vent area of 0.045 m (70 in ) and a cross-sectional area of 25 m2 (270 ft ). Instead of using caustic-water contamination, this test was initiated with anhydrous TEA so as to eliminate water as a possible ingredient in promoting the non-foamy behavior. Out of an initial charge of 54 g, the venting initiated at 8 psig resulted in a loss of 6.4 g material or 12% of the initial mass. Based on an energy balance, the average exit quality can be approximated by

which yields a quality of 0.31. This represents a high quality two-phase flow and is consistent with the churn-turbulent assumption.

As for the last temperature excursion event, no flow regime data was obtained in the VSP. This was made difficult by the fact that the depres-surization technique as employed earlier may not apply because the sample was at 0 bar g and the reaction appeared to be nontempered. The inference from the described test in Figure 6 that vapor/gas flow occurred in a small sample may be very misleading. In view of the high volumetric generation rate, it may be prudent to assume homogeneous vessel venting in this case.

Vent Sizing Methodology

When detailed physical properties and kinetic data are available, integral computer simulations such as DIERS SAFIRE Program [7] can be used to perform the vent sizing calculations. On the other hand, analytical equations have been developed and shown to yield similar results in most instances [8,9]. In addition, VSP data can be applied in a most straightforward manner. Thus these analytical vent sizing equations are employed throughout this study.

Nontempered Reaction. The seemingly nontempered reaction at the end of the tempering period should be treated similar to a gassy system where an adequate relief area would yield a volumetric discharge rate equal to the volumetric gas/vapor generation rate Q T [10], i.e.,

Here the discharge fluid density should be conservatively based on the vessel average density (homogeneous venting) and the mass velocity can be approximated by the Bernoulli equation for low pressure drop situations. Thus Equation (3) becomes

where mt is the mass of the vessel content prior to the last excursion. Based on the measurements presented in Figure 6, the volumetric generation rate is calculated to be (3.2 x 10 kg/min)/(l.l kg/m3 ) or 4.9 x 10 m3 /s for an initial charge of 55 g sample. Scaling up to a 80 m vessel containing 44,400 kg of material, the QT value would be 39 m

3/s. Thus according, to Equation (4), the required vent area for a maximum pressure of 143 kN/m (6 psig) is

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In the above calculation the mass depletion at the onset of the last excursion is taken to be similar to the small-scale result, namely about 45% of the initial mass remains. This large vent size is impractical in practice, thus necessitating the employment of an effective acid quench to prevent such an exotherm in post-relief consideration.

The strategy here is to design an adequately sized relief system to handle the initial polymerization runaway reaction only. With the reaction being tempered in the early phase, the appropriate vent sizing equations are discussed in the next section.

Tempered Reaction. For reference, the vent sizing equation for homogeneous-vessel venting in this category is given by [8]

For churn-turbulent flow regimes, as was found to be applicable here, the asymptotic solution scheme as discussed in Reference [9] is recommended:

where A, ranges from 1.18 to 1.53 and a is the surface tension. Using a a value of 0.0172 N/m for HCN and a A, value of 1.18 (this lower value is more conservative as it results in less vapor-liquid disengagement), the bubble rise velocity according to the above equation is 0.15 m/sec.

Table 3 presents the results for the worst case scenario based on the aforementioned standard contamination mixture. Here the vent sizes for considering churn-turbulent flow (non-foamy) are significantly smaller than the very conservative estimates based on homogeneous venting (foamy regime). As a result of larger degree of vapor-liquid disengagement, the two-phase flow quality entering the vent is substantially higher than the vessel average. For a given area, this has the effect of greatly increasing the volumetric discharge rate, hence much smaller vent sizes are needed relative to the homogeneous venting case. The relief set pressure was found to be an important parameter for consideration. The lower the set pressure, the lower is the self-heat rate or energy release rate. This reduces the corresponding vapor generation rate thus leading to smaller vent sizes. In fact, Table 3 shows that at the lower relief pressure, the vent size for considering churn-turbulent regime comes close to all-vapor venting case.

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Final Remarks. In general, a higher filled vessel requires a larger vent. In addition to a larger total energy release as a result of bigger reactive mass, a higher filled vessel is more conducive to two^phase flow in the relief discharge. With a simulated vent of 0.045 m (70 in ) which more than covers the size range for churn-turbulent regime up to 0.84 bar g (12 psig) relief setting, see Table 3, the amount of liquid entrained is far less than homogeneous venting assumption. The particular test showed that about 8% of the initial mass came off as liquid and 4% as vapor.

Despite the tempering effect provided by HCN vaporization upon relieving to 0 bar g, the later exotherm poses a greater demand for pressure relief. It is therefore prudent to quench the reaction soon after emergency relief. In view of the large amount of water and acid necessary to provide an effective quench, it is recommended that the storage tanks be kept less than 60% full so as to allow enough head space for quenching.

For completeness the vent sizing calculational steps with the thermal data collected are illustrated in the Appendix.

APPENDIX

VENT SIZING ILLUSTRATION

In the region of interest, the P-T data and the self-heat data can be represented by the following expressions

The following physical properties are used:

So far the analysis and experimental data base for treatment of vapor disengagement have been limited to vertical vessels of uniform cross-sectional area. In the current applications, however, the vessels are of horizontal cylindrical geometry (L/D = 3 typical). The representative cross-sectional area can be approximated by the total vessel volume divided by the height (or diameter) of the cylinder, i.e., Ax - V/D, particularly for situations where the vessels are less than 80% full For this sample calculation,

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This criterion allows a quick check as to whether two-phase flow would occur.

Complete Disengagement. Complete vapor disengagement would occur if the two-phase level swell due to the required vaporization rate is not high enough to reach the top of the vessel. For the churn-turbulent regime, the initial vessel free-board volume fraction a has to satisfy the following inequality for complete disengagement,


This is the required relief vent flow in mass per unit time. Next one has to evaluate the mass velocity G which is a function of the exit quality for a given pressure. According to Reference [9], the average exit quality during the overpressure interval for churn regime can be given by

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The two-phase specific volume exiting the vessel is given by

and the critical flow correlating parameter ώ [12] is calculated according to ,2

To check if the flow is choked, we evaluate the critical pressure ratio [12]

which would imply a critical pressure at the choked location of 0.586(157) = 92 kN/m . Since the ambient pressure is normally 101 kN/m , the flow is therefore not choked. The pressure ratio should then be given by

2 (If is used instead, the value is somewhat larger, about 443 kg/m s.) Finany, the required vent area is simply given by

Note that this is an ideal vent area (discharge coefficient of unity). For cases where a vent line of significant L/D is encountered, the actual vent area has to increase to accommodate the reduction in flow due to frictional losses. Reference (13) provides a convenient design chart for estimating the discharge coefficient for two-phase flow.

ACKNOWLEDGEMENT

The HCN work was jointly funded by the following companies: American Cyanamid Company, BASF Chemical Ltd. (UK), Ciba-Geigy Corporation, Degussa Corporation, E. I. DuPont de Nemours & Company, ICI (UK), Monsanto Corporation, Rohm and Haas Company, W. R. Grace Company, and Sohio Chemical Company. The testing facility was provided by American Cyanamid Company at the Fortier Plant. Analyses and insight provided by D. M. Miall of ICI and A. Shaw of Monsanto Corporation are gratefully acknowledged.

SYMBOLS USED 2 = ideal vent area (m )

= constant in bubble rise velocity equation (-) 2

- vessel cross-sectional flow area (m ) - discharge coefficient (-)

A

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- liquid specific heat (J/kg K)

- diameter of horizontal cylindrical vessel (m)

2

- discharge mass velocity, mass flow rate per unit area (kg/m s)

- gravitational constant (9.81 m/s)

- latent heat of vaporization (J/kg)

- dimensionless variables as defined (-)

- horizontal length (m)

- instantaneous mass in vessel (kg)

- initial mass in vessel (kg) 2

- pressure (N/m )

- total volumetric generation rate (m /s)

- heat release rate per unit mass (J/s.kg)

- temperature (K)

- bubble rise velocity (m/s)

- specific volume (m /kg) 3

- internal volume of vessel (m )

- relief vent rate (kg/s)

- quality or mass fraction of vapor (-) 3

- density (kg/m )

- initial vessel void fraction (-)

- temperature rise above set, corresponding to overpressure (°C or K)

- dimensionless turnaround time (-)

- surface tension (N/m)

- critical flow scaling parameter (-)

- critical pressure ratio (-)

ripts

- exiting the vessel

- liquid phase

- difference between vapor phase and liquid phase

Subscripts

- vapor phase

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- initial

- peak or turnaround condition

- set condition

REFERENCES

1. Gause, E. H. and Montgomery, P. D., July, 1960, "Hydrogen Cyanide Stability and Heat of Polymerization", Journal of chemical and Engineering Data, Vol. 5 (3), pp. 351-354.

2. Walker, M. and Eldred, D. N. , 1925, "The Decomposition of Liquid Hydrocyanic Acid", Ind. & Eng. Chem., Vol. 17 (10), p. 1074.

3. Fauske, H. K. and Leung, J. C , August, 1985, "New Experimental Technique for Characterizing Runaway Chemical Reactions", Chem. Engr. Prog., Vol. 81 (8), p. 39.

4. Leung, J. C , Fauske, H. K., and Fisher, H. G., 1986, "Thermal Runaway Reactions in a Low Thermal Inertia Apparatus", Thermochimica Acta. Vol. 104, p. 13.

5. API RP520. 1976, "Recommended Practice for the Design and Installation of Pressure-Relieving Systems in Refineries, Part I - Design, 4th ed., American Petroleum Institute. Washington, D.C.

6. Fisher, H. G., 1986, DIERS Research Program on Emergency Relief Systems", Chem. Engr. Proe.. Vol. 81 (8), p. 33.

7. Grolmes, M. A. and Leung, J. C., August, 1985, "Code Method for Evaluating Integrated Relief Phenomena", Chem. Engr. Prog.. 81 No. 8, pp. 47-52.

8. Leung, J. C., 1986, "Simplified Vent Sizing Equations for Emergency Relief Requirements in Reactors and Storage Vessels", AIChE Journal. Vol. 32 (11), p. 1622.

9. Leung, J. C , 1987, "Overpressure During Emergency Relief Venting in Bubbly and Churn-Turbulent Flow", AIChE Journal. Vol. 33 (6), p. 952.

10. Leung, J. C. and Fauske, H. K., 1987, "Runaway System Characterization and Vent Sizing Based on DIERS Methodology", Plant/Operations Progress. Vol. 6 (2), p. 77.

11. Wallis, G. B., 1969, "One-Dimensional Two-Phase Flow", Chapter 9, McGraw-Hill Book Company. New York.

12. Leung, J. C , 1986, "A Generalized Correlation for One-Component Homogeneous Equilibrium Flashing Flow", AIChE Journal, Vol. 32 (10), p. 1743.

13. Leung, J. C. and Grolmes, M. A., 1987, "The Discharge of Two-Phase Flashing Flow in a Horizontal Duct", AIChE Journal. Vol. 33 (3), p. 524.

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I.CHEM.E. SYMPOSIUM SERIES NO. 110 I.CHEM.E. SYMPOSIUM SERIES NO. 110

TABLE 1 - Amount of Acid Required for Neutralization in an HCN Mixture

The above data were all obtained in the VSP apparatus using similar vent sizes.

TABLE 3 - Vent Sizing Results for Standard Caustic Contamination

*A11 except S02 were 10% water. in water, pure S02 was added as a gas.

TABLE 2 - Flow Regime Summary

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Figure 1 VSP equipment showing top venting configuration Figure 2 External VSP set-up

Figure 3 Temperature and pressure behavior during HCN runaway polymerization

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Figure 5 Delay time and f i n a l exotherm a t 0 bar g and 1 bar g

Figure 4 Self-heat rate and P-T data during runaway

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Figure 6 Mass loss and pressure drop during tempering at 0 bar g

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