Because of the complexity of the GS process flow sheet, there are a number of opportunities for making improvements in the process that, taken together, should increase deuterium production, reduce the number of separate pieces of equipment, improve energy utilization, and reduce costs. U. S. work on improvements in the early 1960s was described by Proctor and Thayer [P4] and has been used in the first Canadian plants. Later improvements patented by Thayer [T3] have been considered for the newer Canadian plants.

This section will describe one flow sheet improvement patented by Babcock [Bl], which

would increase deuterium production by providing supplementary natural water feed to the hot tower. Burgess [B14] describes computer calculations of the increased production that would be possible if additional natural water were fed to the first stage hot tower of one of Savannah River GS units. This section will derive equations for the improved deuterium production obtainable by feeding natural water to one of the stages of the hot tower of the 24-stage example used in Sec. 11.3 of this chapter in the simplified analysis of the process.

The McCabe-Thiele diagram for this process, Fig. 13.28, shows that the deuterium content of the liquid phase flowing down through the hot tower drops to feed level xF between the third and second stages from the bottom of that tower. By feeding additional hot water at rate F’ to the second or first stage, it should therefore be possible to increase deuterium production P at constant H2S circulation rate G, although at the cost of increased tails assay xw, reduced fractional deuterium recovery, and higher heat requirements.

Analysis of the increased deuterium production made possible through use of supple­mentary hot water feed will be made by reference to Fig. 13.35. Here it is assumed that the flow rate of supplementary feed F’ to the top of stage number ns of the hot tower and the product rate P are so adjusted that the deuterium content of water flowing from stage ns + 1 to stage ns equals that of natural water feed xF, to prevent mixing loss at the supplementary

feed point. Primary natural water feed rate F and gas circulation rate G are assumed to remain unchanged, with their ratio at the optimum value G/F = 2.03 found for the previous case with no supplementary feed. Burgess [B14] has shown that G/F should be increased slightly to obtain maximum benefit from supplementary feed, but this refinement has been neglected to simplify the subsequent analysis somewhat.

Under these assumptions, conditions in the cold tower will remain unchanged. Equation (13.122) still relates the product liquid assay xP, natural feed liquid assay xF, and cold-tower gas effluent assay yF. With nc = 16, G/F = 2.03, xP/xF = 4.0, and = 2.32, yF/xF = 0.4558.

Equation (13.170) for the separation performance of the stripping section of the hot tower is obtained by analogy with Eq. (13.125), by substituting xF for xP and ns for nh.

An equation for the separation performance of the enriching section of the hot tower is derived by reference to Fig. 13.35. The deuterium balance for the entire plant above stage і of the hot-tower enriching section (between the dashed lines) is

PxP + (F-P)xi+1 + GyF = Fxf + Gyt (13.173)

At low deuterium content, the equilibrium relation for stage і is

Fi = XF — (13-174)

ah

By eliminating yt from (13.173) and (13.174), difference equation (13.175) for the liquid-phase deuterium atom fraction x is obtained:

(13.175)

The solution of (13.175) for the boundary conditions xt = xP at і = nh + 1 and xt = xF at ;’ = + 1 is

[P(xP — xF) + GyF — GxF/ah [G/(F — P)ah ] nh ~nS + fxF — GyF
F — G/ah

Equation (13.176) provides an implicit relation between the product/feed ratio P/F and the number of enriching stages nh—ns. For the present case, with xP/xF = 4, yF/xF = 0.4558, ah = 1.80, G/F = 2.03, and nh = 24.

To complete the analysis, it is necessary to find the amount of supplementary feed for the hot tower, F for a given number of stripping stages ns. To do this, W and xw are eliminated from (13.170) by means of (13.171) and (13.172), and the resulting equation is solved for ns + l:

‘ (P/F)(xp/xf -1) + (G/F)(yF/xF — І/a,) 1 / Г G/F

(1 +F’fF-P/F)(ahyF/xF)-( +F’/F-Pxp/Fxf) ]/ + F’/F-P/F)

(13.178)

Substitution of the given values for xP/xF = 4, G/F =2.03, )>f/xf = 0.4558, and ah = 1.80 yields

Numerical solution of Eqs. (13.177) for P/F and (13.179) for F’/F with auxiliary feed to the top of the second stage of the hot tower (ns = 2) or to the first stage (ns =1) yields the results of Table 13.25, where they are compared with the case of no supplementary feed.

Figure 13.36 compares the above results for this 24-plate case, without reoptimization of the feed rate to the cold tower, with Burgess’ [B14] calculations for the Savannah River plant, in which feed to the cold tower was reoptimized for maximum production.

Determination of the economic proportion of supplementary feed to the hot tower involves balancing the advantage of increased production against the extra costs of preheating additional feed water and stripping H2S from additional waste. In the stripping section of the hot tower, larger downcomers would be needed for the increased liquid flow, and at some value, a larger tower diameter. In a new plant designed for it, some supplementary feed to the hot tower would seem to be advantageous. It would probably be neither practical nor economical to use more than 50 percent extra feed to the hot tower.