Worldwide Catalyst Report Three-step procedure optimizes FCC slide-valve differential pressure

Ray Fletcher
Akzo-Nobel Chemicals Inc.
Houston

Additional catalyst circulation rate often directionally improves product selectivity by maximizing the number of active sites the oil encounters while passing through the riser. This improvement, in turn, increases unit profitability.

Slide-valve differential pressure can be optimized using a logical, three-step approach. This procedure involves optimizing:

• The aeration rates in the standpipe

• The pressure balance between the reactor and regenerator vessels

• The physical properties of the catalyst.

Standpipe aeration

Slide valve limitations often are the result of less-than-optimal aeration rates in standpipes. Poor aeration rates may result in loss of differential pressure, erratic differential pressures across the slide valve, or, in extreme cases, physical bouncing or hopping of the standpipe.

Careful analysis and adjustment of these rates will, in many cases, increase the slide-valve differential pressure or stabilize operations. For a unit limited by slide-valve differential pressure, an improvement in slide-valve position will allow increased catalyst circulation rates.

The catalyst flux rate in the standpipe first should be calculated and compared to the design flux for the unit in question. Typical flux rates are 200-250 lb/sq ft-sec. If the standpipe is operating well beyond the design flux rate, little improvement may be expected by optimizing the aeration rates.

A column of catalyst in a standpipe builds pressure in the same way a column of water does. The pressure head builds as the depth increases. As the pressure builds, the aeration media surrounding the catalyst particles become compressed, allowing greater contact between adjacent particles.

The purpose of the aeration media injected into the standpipe is to counteract the compression that catalyst and fluidizing media undergo as they pass down the standpipe. If aeration media were not injected into a standpipe, the pressure would build to the point that the interstitial gas would no longer separate the particles, and the standpipe would defluidize.

The minimum fluidization rate is determined by carefully adding fluidization media to a packed column such that the last incremental volume added fluidizes the column. This minimum rate is defined as the incipient fluidization point.

The bed density at incipient fluidization is approximately equal to the apparent bulk density (ABD) as noted on most equilibrium catalyst data reports.

The maximum aeration rate is determined by adding additional aeration media to a fluidized column until the last incremental volume added produces in the column a permanent bubble that becomes an obstruction to the flow of catalyst.

Once this bubble is formed, the catalyst must flow around the bubble. The bubble then acts as an obstruction, limiting catalyst circulation rates in the same way a shovel blade or spelled lining obstructs catalyst flow. This point is defined as the incipient bubbling point.

If the aeration rates at any of the taps along the length of the standpipe are not set correctly, the result could be a loss of head as measured at the inlet to the slide valve. This may be caused by too little or too much aeration media.

The best method to evaluate whether standpipe aeration rates have been set properly is to perform a pressure survey using the same pressure gauge at each aeration tap along the standpipe. A steady increase in pressure should be observed as the depth of catalyst increases (Fig. 1 [25300 bytes]).

Improper aeration rates will result in an inflection point at one or more of the taps (Fig. 2 [25464 bytes]). It is important to note that the inflection point occurs at the aeration tap that has an improper aeration rate. The pressure stops building at this point and, in severe cases, a loss in pressure may be observed.

An inflection point is corrected by optimizing the aeration rate at that point. This should result in increased pressure at the inlet of the slide valve and improved slide-valve differential.

The first step in optimizing the aeration rate involves calculating the amount of compression the catalyst and fluidizing media undergo as they pass down the standpipe. The calculation must begin at the inlet to the standpipe and progress from one aeration tap to the next.

The volume contraction of the aeration media caused by the increase in pressure establishes the theoretical amount of aeration media required at each tap. In practice, adding the theoretical required aeration in most cases will result in reaching or exceeding the incipient bubble point. Most units require 50-75% of the theoretical volume.

A simple method for calculating required aeration rates is presented in Fig. 3 [9449 bytes].^{1 2}

Pressure balance

The pressure of the conversion section of the FCC is set primarily by two variables (Fig. 4 [17696 bytes]):

• The set point for the flue gas slide-valve pressure

• The suction pressure of the wet gas compressor.

Optimization of the pressure balance requires that additional capacity exist on either the combustion air blower or the wet gas compressor.

Regenerator

The discharge pressure of the combustion air blower is the driving force for the pressure balance in the regenerator vessel. This pressure is then controlled by the flue gas slide valve.

An increase in the flue gas slide-valve position increases the flow rate of combustion gasses out of the regenerator, and therefore has the effect of decreasing the regenerator pressure. This slide valve often is used to maintain a minimum differential pressure between the reactor and regenerator vessels.

The pressure exerted by the depth of catalyst in the regenerator dense bed, plus the column of catalyst present in the regenerated catalyst standpipe, combine to establish the pressure observed on the upstream side of the regenerator slide valve. Increasing the pressure set point on the flue gas slide valve or increasing the bed level in the regenerator will increase the pressure on the upstream side of the regenerator slide valve.

Excess main air-blower capacity must exist before an increase in the set point for the flue gas slide-valve pressure is considered. This is because, in order to inject air, the combustion air blower must overcome the pressure increase in the regenerator. An increase in regenerator pressure requires additional combustion air-blower horsepower.

An increase in the set point for the flue gas slide-valve pressure also will result in a loss of reactor slide-valve differential pressure. An increase in the reactor slide-valve position will be required to maintain a constant flow of spent catalyst from the reactor into the regenerator after the regenerator pressure increases.

An increase in regenerator slide-valve position is therefore gained at the expense of the reactor slide-valve position. In other words, an increase in the regenerator pressure trades reactor slide-valve position for regenerator slide-valve position.

An increase in regenerator bed level must be approached with caution because of the potential for increased catalyst losses out of the regenerator. Losses may increase as the bed level is raised. The increase in losses can be significant if the minimum transport disengaging height of the regenerator vessel is exceeded.

In this case, catalyst traffic into the cyclones will increase substantially because of insufficient catalyst disengaging height. Catalyst losses also will increase considerably if the cyclones are operating close to their maximum design rate before the dense bed level is increased.

Reactor

The suction pressure of the wet gas compressor controls the reactor pressure balance from the base of the riser, through the main fractionator, and into the fractionator overhead cooling system. A decrease in the suction pressure of the wet gas compressor, therefore, will lower the pressure downstream of the regenerated-catalyst slide valve.

The pressure on the upstream side of this slide valve will remain fairly constant. The effect of a drop in the suction pressure of the wet gas compressor, therefore, is an increase in the regenerator slide-valve differential pressure.

Decreasing the wet gas suction pressure increases the actual volume of wet gas the compressor must compress as a result of the volume expansion of this stream. Excess compressor horsepower is therefore required before this change can be made.

A decrease in the suction pressure of the wet gas compressor decreases the pressure on the upstream side of the reactor slide valve. This slide valve is then required to increase its valve position in order to deliver a constant flow of spent catalyst into the regenerator. Regenerator slide-valve differential pressure is therefore increased at the expense of the reactor slide valve.

This results in a loss of reactor slide-valve differential pressure. Decreasing the suction pressure of the wet gas compressor therefore trades reactor slide-valve differential pressure for regenerator slide-valve differential pressure.

If the regenerator slide-valve differential pressure is limiting the catalyst circulation rate and excess capacity exists on either the combustion air blower or the wet gas compressor, an opportunity to optimize the slide-valve position and differential pressure exists.

Increasing regenerator pressure or decreasing the suction pressure of the wet gas compressor will result in reduced reactor slide-valve differential pressure, with a corresponding increase in the regenerator slide-valve differential pressure. Either control move will result in increased differential pressure on the regenerator slide valve and will enable the unit to circulate more catalyst.

Catalyst properties

In many cases, optimization of catalyst physical properties will result in improved fluidization characteristics and increased slide-valve differential pressures. This is especially critical in units with unusual standpipe configurations such as sloped standpipes, standpipes with bends or doglegs, and standpipes with large distances between aeration taps.

Optimization of catalyst physical properties has resulted in improved circulation and slide-valve differentials on several FCC units. One useful approach to determining the fluidization characteristics of any catalyst system is the method pioneered by Abrahamsen and Geldart.³ Their work focuses on the gas velocity of the catalyst and fluidizing media in the standpipe at incipient fluidization and bubbling.

This approach allows the process engineer to characterize and maximize the range of acceptable aeration rates within a given standpipe. This range of stable operations is critical on units with unusual standpipe configurations, which allow excessive compression of catalyst and fluidizing media. Excessive compression of the catalyst and fluidizing media will approach a packed-bed situation and must be avoided.

Several useful observations may be derived from this work.

The unit engineer can use these observations to optimize fluidization characteristics on the FCC. The first two equations to consider are those describing incipient fluidization and bubbling (Equations and Nomenclature).

The ratio of the gas velocity at incipient bubbling over the gas velocity at incipient fluidization is known as the Maximum Stable Expansion Ratio (MSER). This ratio is very useful in defining the relative range of stable operations in a given standpipe (Equation 3).

The ratio in Equation 3 defines the relative range of fluidization or the "size of the window" of stable operations for a given catalyst system. The larger the ratio, the more likely a particular catalyst will fluidize over a wide range of operating conditions with a fixed aeration rate.

Careful attention to this ratio, combined with optimization of the catalyst physical properties, may result in improved standpipe operations for units with abnormal standpipe configurations or units that are limited in slide-valve differential pressure.

The key points to recognize in Equation 3 are:

• The weight fraction of 0-45 µ particles is present in an exponential term in the numerator of the characterization ratio.

  This indicates that the greater the percentage of fines in the circulating catalyst inventory, the more likely that system will fluidize. The magnitude of this ratio increases dramatically as the amount of fines increases. With the increase in the characterization ratio comes an effective increase in the window of stable operations available in a given standpipe. This weight fraction will, in many cases, override all other catalyst physical properties.

• The mean particle diameter is found in the denominator of this equation.

  This implies that as the mean particle size of the catalyst in the circulating inventory decreases, the ability to circulate catalyst improves. The denominator of the equation is reduced, which increases the magnitude of the ratio. As a result, the standpipe is more likely to remain fluidized over a wide range of operations.

• The particle skeletal density is found in the denominator of the ratio.

  This clearly implies that as the particle density decreases, the range of stable standpipe operations will increase. This is because a larger expansion ratio results from a reduction in the magnitude of the denominator. A lower-density catalyst, therefore, will be desirable in units experiencing difficulty in fluidizing catalyst.

The ability to increase slide-valve differential pressure or to circulate additional catalyst will be enhanced by increasing the 0-38 µ fraction of the circulating inventory, and by using a catalyst with lower skeletal density. These two moves must be made with caution in units not equipped with electrostatic precipitators or flue gas scrubbers, and in units that have tight environmental limits on particulate emissions.

An increase in the fresh catalyst activity also will tend to move a unit limited by catalyst circulation away from a slide-valve limitation. This is because an increase in regenerator temperature reduces the catalyst circulation rate.

References

1. Kuneii, D. and Levenspiel, 0., Fluidization Engineering, 2nd ed., Butterworth-Heinnemann, Newton, Mass., 1991, pp. 376-81.

2. Mott, Raymond W., Katalagram No. 83, 1992, W. R. Grace & Co., Davison Chemical Division, Baltimore.

3. A.R. Abrahamsen and D. Geldart, Powder Technology, Vol. 26, p. 39-40, 1979.

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