reducing the purge in flare stacks

Reducing the purge in hydrocarbon vent stacks

Stephen G. Bryce and Robert E. J. Fryer-Taylor Shell Research Limited, Thornton Research Centre, PO Box 1, Chester CH13SH, UK

The purging of hydrocarbon vent stacks is necessary to ensure that large flammable regions are not created within such facilities. In order to minimize the environmental impact of purging, it is vital that purge rates should be set at the lowest level consistent with safe operations. The methods which are currently available for determining these minimum rates are empirical and may be overly conservative; therefore, it is important to develop a more robust basis on which to calculate the required gas flows. To this end, a reduced-scale Rayleigh-scattering investigation of purging flows has yielded two-dimensional concentration maps for a range of purge rates and crosswind velocities. The results indicate that it may be possible to reduce the purge rate significantly without adversely affecting operational safety; full-scale testing must be performed to determine the potential hazards associated with such action.

Keywords: vent s t ac k s; purg ing; Ray le igh sc at t er ing

The venting of gaseous hydrocarbons may be required to maintain safe conditions within a production or processing facility. During the venting process, the momentum of the gas flow limits the ingress of air to a region within a few diameters of the tip of the stack. However, when the stack is not venting, it is possible for the entire stack to become filled with air; on resumption of venting, potentially flammable regions could be created throughout the stack. There is only limited quantification of the overpressures that would be generated by igniting such regions within a complex vent system’. It is possible that these pressures could be sustainable by the structure but this has not been established. Therefore, it is vital to prevent large-scale ingress of air into a non-venting stack.

The first problem may be minimized by high purge rates, but this wastes gas and exacerbates the second problem. At present, purge rates are determined on the basis of the criteria of Husa* and Tan’, but these are empirical correlations and may be conservative in their recommendations. Therefore, the objective of the present investigation is to determine a definitive quantified criterion for the lowest possible safe purge rate within a vent stack.

Experimental scaling

To meet this requirement, it is usual to maintain a ‘purging’ gas flow through the stack under nonventing conditions. A range of purge gases has previously been studied2, and predictions have been made of the minimum safe purge rates for these gases. It was found that the required purge rates decreased with increasing molecular weight of the purge gas. However, it is most feasible to purge a stack using the gas that is being produced, which in the majority of cases leads to the use of a hydrocarbon mixture. There are two major problems associated with the use of gaseous hydrocarbons:

1. The purge gas is flammable, so the possibility remains of flammable regions existing within the stack.

2. Gases such as methane have been identified as ‘greenhouse gases’; these may potentially contribute to a ‘global warming’ phenomenon.

A simplified representation of a purge flow is shown in Figure I: the flow will be driven by buoyancy forces, inertia forces and viscous forces. In the absence of any large-scale heat-transfer effects, the relevant non- dimensional parameters are

Reynolds number: U,dlv Richardson number: (pO - p,)gd/(p,Ulz) buoyancy ratio: velocity ratio: EiYrn E 1

where U, is the velocity of the purge gas, U, is the crossflow velocity of the air, d is the diameter of the stack, Y is the kinematic viscosity of the purge gas, g is gravitational acceleration, po is the density of air and p1 is the density of the purge gas.

It has been observed3*4 that the purge gas velocity required to maintain a constant oxygen level at a fixed depth within a stack is proportional to the stack diameter, i.e. maintaining a constant Reynolds number between stacks will not maintain a constant gas concentration distribution. Preliminary experiments

095&4230/94/030249-07 0 1994 Butterworth-Heinemann Ltd

J. Loss Prev. Process Ind., 1994, Volume 7, Number 3 249

1 Flammable region

t

Air

Purge gas

Figure 1 Simplified purging flow

also indicated that Reynolds scaling was not appropriate to this investigation, and so only the Richardson number, buoyancy ratio and velocity ratio were matched at experimental and full scale. This matching was achieved for l/12, l/15 and 112.5 scale models of a 0.254 m diameter stack. A natural gas purge discharging upwards was simulated by the discharge downwards of a Freon/air mixture for a range of purge rates and crosswind velocities.

Concentration measurement by a Rayleigh-scattering technique Previous investigation2 of the purging of vent stacks has relied on gas sampling devices for the measurement of gas concentrations (either purge gas or oxygen) at the centreline of stacks. While such devices can provide a history of concentration, they can only do so at a relatively small number of points within the stack. It was determined that a field measurement technique, i.e. one which could provide simultaneous measurements across a two-dimensional region of the flow, would greatly facilitate the investigation.

If a gas flow is seeded with a light-scattering tracer of consistent quality, the light-scattering intensities (the so-called ‘Rayleigh scattering’) may be directly related to local gas concentrations. This technique was used in this investigation.

Reducing the purge in hydrocarbon vent stacks: S. G. Bryce and R. E. J. Fryer-Taylor

Figure 2 is a simplified schematic representation of the experimental system. A laser sheet illuminated a two-dimensional vertical section in the centreline plane of a quartz-walled stack of circular cross-section. The smoke-seeded mixture of Freon and air flowed downwards within the stack towards the open end, and the resultant light scattering was video-recorded via a CCD camera. A fan could be used to provide a controllable wind across the open end of the stack, where the wind velocity was measured using a pitot- static tube.

The basic video-recorded image provided quali- tative information on gas concentrations within the stack, but further processing was required to obtain quantitative measurements. This processing corrected for variations in the background signal and provided a reference image for 100% gas concentration throughout the stack. A simple division process then yielded a quantitative image, which could be false-coloured to provide a concentration map for the illuminated region of the stack. These maps, which could be determined for both the time-averaged and time-varying behaviour of the flow, formed the basis for measurement of flammable zones.

The effects of purge velocity at zero crosswind The effects of purge velocity at zero crosswind are typified by the results for the l/12 scale stack. The mixing processes within the stack can be determined by reference to Figure 3 which shows a single-frame image of the ten diameters of the stack closest to the tip for a model-scale purge velocity U,, = 4.4 mm s-’ (full-scale lJ, = 15.8 mm s-l). It can be seen that the mixing process between the purge gas and air was highly vertical, where the magnitude of the vorticity was enhanced by the close proximity of the stack wall. This vorticity was a consequence of the density

Gas + smoke

Quartz-walled

.. stack

” . ” . . . .

‘. .. ‘ . ..,.

CCD camera ” . . . . . .

‘ . ._

B -2

Figure 2 Schematic representation of the experimental system

250 J. Loss Prev. Process Ind., 1994, Volume 7, Number 3

Figure3 Single-frame image: U,, = 4.4 mm s-’ d,,, = 0.021 m

difference between the gases and could be most clearly visualized at the lower purge rates. As the purge rate

was increased, the vigour of the mixing increased as the vortices within the stack reduced in size but increased in rotational speed.

A time-averaged image of Figure 3 is shown in Figure 4. The flow was averaged over 32 video frames (1.28 s) and it can be seen that most of the time- dependent behaviour has been removed. It was decided that, whilst such a time-averaged method would not resolve the unsteady nature of purging flowfields, it would meet the stated objectives of the experimental programme for practical purposes. Therefore, the results presented in the remainder of this paper are based on time-averaged processing.

Figure 5 highlights the flammable regions on concentration maps for the stack at eight different purge rates from 1.1 mm s-l to 8.6 mm s-l (U, = 4.1 mm s-l to 30.9 mm s-l). It can be seen that, at lower purge rates, the vertical nature of the mixing between the purge gas and air resulted in a flammable region within the stack. This flammable region comprised several distinct flammable ‘pockets’, and these did not always encompass the stack centreline. It is therefore misleading to consider only the centreline concentration profile, such as shown in Figure 6, when determining the length of the flammable zone. In this investigation, the flammability hazard was quantified as that percentage of the illuminated

Figure 4 Time-averaged image: U,, = 4.4 mm s-‘, d.,, = 0.021 m

region of the stack where the air concentration lay between 85 and 95%. This quantification was accomplished by extracting the histogram of concentrations for each case, a typical example of which is shown in Figure 7. The flammable volume was then divided by the volume of the illuminated region and multiplied by 10 to yield an effective flammable length in stack diameters.

The effect of purge velocity on the flammable volume can be seen in Figure 8: an increase in purge velocity generally led to a reduction in the effective flammable length, except in the mid-velocity range around U ,,,, = 3.2 mm sP’. At this intermediate purge velocity, there was an apparent anomaly in the decreasing trend; however, inspection of every other case revealed that this was a real effect. At some value of purge velocity, the effect of an increase was to maintain or increase the overall extent of the flammable region, whilst further increases again led to a reduction in the extent of this region. It is hypothesized that this effect is a consequence of a balance between vortex strength and purge velocity, where an increase in vortex strength results in a partial blocking effect on the discharging gas. At lower values of purge velocity, the vortex is not strong enough to have this effect, whilst the momentum of the flow at higher values can overcome the blockage.

Despite this effect, it was concluded from the studies at the three scales that, for a zero crosswind, a full-



Figure 5 Flammable regions for d,,, = 0.021m: U, = 1.1, 1.8, 2.5, 3.2, 3.8, 4.4, 7.0, 8.6mms-’

50 Ml 70 80 90 Air concentration, %

Figure 6 Centreline concentration distributions: d,,, = 0.021 m. u = Irn 1.8 mm s-’

scale purge velocity Ur = 8 mm s-r would ensure that the flammable mixture within the top ten diameters of the stack tip would occupy less than 50% of the available volume, i.e. the effective flammable length would be less than five stack diameters.

The effects of crosswind The effects of crosswind velocity on gas concentrations within the stack are discussed with reference to the results for the l/12 scale model, although the same general trends were identified at all scales. For this

0 0 10 20 30 40 50 60 70 80 90 1

Air concentmion sb

Figure7 Histogram of air concentrations: d,,, = 0.021 m, U,, = 1.8 mm s-l

stack, the crosswind range investigated was from u =3.0 m s-l to 9.6 m s-r, corresponding to a rat;, of U, = 20-65 knots at full scale.

A previous study indicated that a crosswind would decrease the length of any flammable zone (as measured along the stack centreline) whilst moving the zone further within the stack. As can be seen from Figure 9, the same trend could be identified in these tests at low crosswind velocities, but at higher crosswinds this centreline flammable length was increased. In most cases where the off-centreline behaviour was included, the overall extent of the flammable region was increased even at low values of crosswind velocity.


Indeed, these results, which are plotted in Figure 10, show that, at higher values of purge rate and crosswind velocity, a flammable zone may be created where previously it did not exist. The intermediate purge rate phenomenon, referred to in the preceding section,

was again present. A consistent effect at all purge rates and crosswind velocities was to increase the air concentration in the exit plane of the stack above that for the equivalent zero crosswind case.

Purge velocity mm/s

uwm = 6.8 m/s

It is believed that the increase in flammable extent with crosswind is a function of flow separation along the wall of the stack close to the tip. This behaviour is illustrated in Figure 12 which clearly shows the generation of small-scale vortices following separation. This vertical behaviour had a marked effect on concentration in the region as the vortices generated considerable mixing along the wall. This flowfield is schematically represented in Figure 12, where the primary effect of the crosswind is the creation of a large-scale vortex close to the stack tip. Such a flow phenomenon is well-known for a crossflow above a slot in a surface. The low pressure core of the vortex generates a negative pressure gradient, both along the axis of the stack towards the tip and radially towards the axis. Consequently, the flow along the walls of the stack will separate. More air is entrained into the low pressure area at the tip of the stack, and this is conveyed downwards along the wall of the stack within the separated region. As a result, air is present further within the stack than for a zero crosswind case. Along the axis of the stack, the gas flow is accelerated towards the tip by the negative pressure gradient. Therefore, the centreline gas flow mixes more vigor- ously with the air at a greater distance from the tip, and the result on the centreline may be a shorter flammable length further from the tip. However, the increased mixing within the separated region along the wall ensures that the overall extent of the flammable region is increased.

Figure8 Variation of effective flammable length with purge velocity: d,,, = 0.021 m


70 74 78 82 86 90 94 98 Air concenmarion B

When these off-centreline effects were included, it was determined from the experiments at the three scales that, for crosswinds up to 65 knots, a full-scale purge velocity of 13 mm s-l would ensure that the effective flammable length would be less than five stack diameters.

Figure 9 Effect of crosswind on centreline concentration distribution: U,, = 1.8 mm SC’, d,,, = 0.021 m

8.0

3 7.0 uw= 10.7rds 1 . uw= 16.4rds 6.0

A Uw=24.4ds x Uw=34.5ds

Q 5.0 4 1! 4.0 D

I 3.0 E .E 2.0 j 5 1.0

0 0.004 0.008 0.012 0.016 0.02 0.024 0.028 0.032

Full-scale purge velociry m/s

Figure 10 Variation of effective flammable length with purge rate and crosswind velocity Figure 11 Enhanced image of crosswind-induced separation



Crosswind

Purge

seen that the l/25 scale results indicate that a full- scale purge velocity of 8 mm s-l would be required to ensure that the equivalent flammable length was less than five stack diameters, whilst, at the l/12 scale, the required value to achieve this flammable length would be 5 mm s-l, i.e. a reduction of 37.5% in the gas requirement.

Whether this behaviour was a result of a scale effect or an experimental anomaly can only be determined by full-scale validation testing. In the absence of full-scale information, it must be assumed that the higher value of purge velocity is correct. The impact of this requirement on predictions of minimum purge rates is considerable.

A new criterion For minimum purge rate Laboratory-scale experimentation has highlighted the need for full-scale tests to validate the measurements of flammable volumes at very low purge rates. These tests would also validate the scaling criterion that has been employed in this study; if its suitability is confirmed, this criterion will greatly facilitate the prediction of minimum safe purge rates.

Figure 12 Schematic representation of crosswind effect

The effects of experimental scale The effects of experimental scale could be most clearly determined from the results at zero crosswind. Figure 13 shows the variation of effective flammable length with purge rate for the l/12, l/15 and l/25 scale investigations, where the purge velocities have been scaled to their full-scale values for a 10 inch diameter stack. It was found that there was very good agreement at all three scales for zero crosswind above a full-scale purge velocity of 9 mm s ‘. Below Ur = 9 mm s-r, there was a notable difference between the results at l/12 scale and those at l/15 and l/25 scale. It can be

It has been found from the reduced-scale experimental programme that any minimum purge criteria based on either measurements or predictions of centreline concentrations2,3 cannot recognize those large flammable volumes that do not encompass the stack centreline: these criteria are therefore unsuitable. Therefore, it is proposed that the minimum purge criteria for a stack should be based on the maximum permissible effective flammable length within the vent system; the quantification of this maximum flammable length will also necessitate full-scale testing.

In order to achieve this quantification, it will be necessary to measure the overpressures generated within a vent system by ignition of a flammable mixture for progressively more severe scenarios. These scenarios would range from ignition of a small flammable pocket close to the stack tip to ignition of a fuel-rich flammable mixture filling the complete vent system’. The maximum permitted flammable length could then be established by comparison of overpressures within the system with the limiting stress for its structure.

. 1/25th scale + 1/15th scale 4 l/lZth scale

It is possible that such a study would indicate that the worst-case overpressure would not damage the structure for particular system geometries. In such cases, it would then be possible to employ a zero purge rate, i.e. the purging requirement would be removed. Where structurally damaging overpressures were possible, analysis of the concentration measurements from both laboratory and full-scale tests would indicate the purge rates required to ensure that flammable lengths did not exceed their specified limiting values.

2 6 10 14 18 22 26 30 34 Full-scale purge velocay (mm/s)

Figure 13 Variation of flammable length with purge rate and experimental scale

If the full-scale concentration measurements con- firm the dominant role of the Richardson number for this type of flow, it will be possible to express a minimum purge rate criterion in terms of the critical


Reducing the purge in hydrocarbon vent stacks: S. G. Boyce and R. E. J. Fryer-Taylor

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8 B 9

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diameters and if the l/12 scale data are validated, Figure 3 shows that Ut, = 4 mm s-l for a zero crosswind. Consequently, the criterion would alter to

u,, = 0.0079d”-

which is also plotted in Figure I4 for comparison. If validated, this result would enable large reductions in purge rate.

0.2 0.4 0.6 Stack diameter m

It can be seen from the results of Figure 14 that slight increases in permitted flammable length and experimental scale have a considerable effect on the predicted minimum purge rate criteria. This behaviour emphasizes that full-scale studies will play a major role in achieving the maximum possible reductions in purge rates.

Figure 14 Comparison of minimum purge rate criteria Conclusions

Richardson number (Ri,) above which the maximum permissible flammable length would be exceeded. A sensitivity analysis has been performed in order to determine the effect on such a criterion of variations in permitted flammable length and of the ‘scaleability’ of the reduced-scale data.

For example, if it is found that this maximum length is five diameters (which is likely to be a very conservative result) and the l/15 scale data of Figure 13 are validated, for zero crosswind conditions the full-scale purge velocity should be no less than 8 mm s-l i e. U rc = 8 mm s-l. This velocity requirement corresponds to Ri, = 27900 for a methane purge, and, consequently, Ri must be less than 27900 for any stack. This criterion may be expressed as

u,, = 0.0159do.5

and is plotted in Figure Z4 for stack diameters up to 0.52 m. It can be seen that this result is in very close agreement with the current industry practice based on Husa’s criterion2. Such a result would not permit any reductions in purge rate, and therefore would be disappointing.

However, if the overpressure data indicate that the maximum permissible flammable length is six

A reduced-scale experimental investigation has quantified the effects of purge rate and crosswind velocity on gas concentrations within hydrocarbon vent stacks. It has been shown that increasing the purge rate at zero crosswind will generally decrease the flammable volume within the stack. It was also found that the main effect of crosswind was to increase the flammable volume within the stack through enhancing mixing. Indeed, under strong crosswind conditions, increasing the purge rate may create a flammable volume within a stack which was previously fuel-rich. Overall, the results indicate that it may be possible to significantly reduce the purge rate utilized in hydrocarbon vent stacks. However, it will first be necessary to resolve uncertainties about scale effects on the flowfield at low purge rates, and to quantify the safe Bammable length for vent systems.

References 1 Tite, _I. P., Greening, K. and Sutton, P. Chem. Eng. Res. Design

1989, 67 2 Husa, H. W. Hydrocarbon Processing and Petroleum Refiner

1964, 43(5) 3 Tan, S. W. Hydrocarbon Processing 1967, 46(l) 4 G.K.N. Binvelco Ltd. ‘Design of offshore flare systems and cold

vents - background document’, Offshore Technology Report OTH 89 305, 1990


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