A new process for HNO3 recovery from spent pickling solutions

Resources, Conservation and Recycling, 9 ( 1993 ) 141-152 141 Elsevier Science Publishers B.V.

A new process for HNO3 recovery from spent pickling solutions

Guido Saracco and Maurizio Onofrio Dpt. Scienza dei Materiali ed lngegneria Chimica, Politecnico di Torino, Italy

(Accepted 18 January 1993)

ABSTRACT

During the proposed process the spent pickling bath is at first neutralized with powdered lime, which causes the precipitation of metal hydroxides and CaF2. After concentrating the neutralized solution, an acidification by addition of H2SO4 is performed in order to regenerate HNO3 as CaSO4" 2H20 precipitates.

The effectiveness of the process is examined for its chemical, technical, and economical characteristics.

INTRODUCTION

Nitric-hydrofluoric pickling of stainless steel generates spent acid solutions whose environmental impact is hard to control mainly because of their high content of nitrates. The authors consider that the best answer to this problem is the recovery of NO~- as HNO3 to be reutilized in the pickling tanks.

Several processes have already been proposed for this purpose, based either on merely chemical means [ 1,2 ], or on ionic-membrane technology [3,4,5 ]. As already briefly reported elsewhere, [6], we have proposed and tested a new process whose main features are shown in Fig. 1.

EXPERIMENTAL ANALYSIS OF THE PROCESS

Spent pickling solutions were collected from ILVA-Divisione Inox (Torino, Italy) which financially supported this work.

The composition of the baths is strictly linked to the type of steel, the thick-

Correspondence to: G. Saracco, Dpt. Scienza dei Materiali ed Ingegneria Chimica, Politecnico di Torino, C.so Duca degli Abruzzi 24, l 0129 Torino - Italy.

0921-3449/93/$06.00 © 1993 Elsevier Science Publishers B.V. All rights reserved.

142 G. SARACCO AND M. ONOFRIO

1

NEUTRA- LIZATION

t SLUDGE

SEPARATION AND WASHING

I 1

t RECOVERED 1 I 1 NITRIC ACID 1

Fig. 1. Flow-diagram of the proposed process.

ness of the sheets, the annealing temperature, etc. We tested the two spent baths which contained the maximum and the minimum content of nitrates:

(A) NO,: 73.6 g/l; F- : 53 g/l; Fe3+: 39.6 g/l; Cr3+: 6.9 g/l; Ni2+: 5.6 g/l;

(B) NO, : 37.1 g/l; F-: 40 g/l; Fe3+: 34.1 g/l; Cr3+: 6.5 g/l; Ni2+: 4.9 g/l. Each step of the process has been experimentally investigated. The results

are described and discussed in the following paragraphs.

The neutralization The use of powdered CaO as a neutralizing reagent, instead of milk of lime,

greatly reduces the following evaporation step by keeping nitrates relatively concentrated. Overall kinetics are thus governed by the dissolution of lime particles, which tend to get covered with reaction products whose synthesis is faster than the dissolution itself. It is therefore important to study the influ- ence of different excesses of reagent beyond the stoichiometric dose ( 117 g/l of 9 1% (technical efliciency ) CaO for bath A and 74 g/l for bath B) on the speed in reaching the desired basicity ( pH 9 ) .

During each test the stated dose of lime was continuously and homogene- ously added during the first 30 min of reaction upon the surface of a l-1 sam- ple of bath by means of a vibrating sieve. This relatively low speed of addition was chosen because the covering of lime particles is intense at the beginning, due to the massive precipitation of Fe( OH), and CaF,, whereas the lime added afterwards reacts more efficiently. Complete suspension was guaran- teed by mechanical agitation during every test, which was repeated three times, getting the mean results shown in Fig. 2 (referring to bath A).

PROCESS FOR HNO3 RECOVERY FROM SPENT PICKLING SOLUTIONS 143

pH

12

10

I Excess

) ~CaO 20% o,O

20 40 60 I I

0 80 100 120 Time (mini

Fig. 2. Variation of the pH value of bath A during neutralization with different excesses of CaO.

After 45 min (a reaction t ime characteristic of a reasonably small reactor size) a 10% excess of CaO was found to be sufficient to neutralize bath A. On the other hand bath B required a 40% excess for the same reaction time. This can be easily explained considering the different thermal evolution of the two baths during neutralization in nearly adiabatic conditions. In both cases the temperature rises during the first 30 min of reaction, due to the heat of hydra- tion o fCaO (AH ° = - 15 kcal/mol, [7] ), up to a max imum value (82°C for bath A and 54 °C for bath B ); afterwards it slowly decreases because of evaporation (2-3% of the reacting mass).

We believe that these temperature changes play a primary role in acceler- ating the dissolution of lime particles.

Moreover, as shown in Fig. 3, a fair reduction of the excess of CaO is achievable by dosing CaO through thinner and thinner sieves. The increase in CaO activity is less than proportional to the rise of the specific surface of the solid particles mainly because of agglomeration phenomena.

Using a 37-ltm sieve instead o f a 105-/zm one (used for all the other tests) a 13% excess of CaO was saved, enabling moreover a small reduction of the volume of left over sludges. These undoubted advantages must be balanced against the cost of performing such a deep sieving at the industrial scale. The use of intensive mixing systems capable of breaking down the lime particles in the reacting suspension was found to be even more expensive due to large energy consumption.

On the other hand, a good way of reduction of the volume of outcoming sludges, whose disposal involves the highest variable cost of the process (see Table 1 ), lies in the recovery of nickel as hydroxide.

Fig. 4 shows the gradual decrease of the concentrations of metallic ions


12 pl-I 11

10

O

I

1

6

5 -

4

2 0 10 20 30

maximum'particle' size (JJl~l) I

/ / /

Excess of CaO I%1

Fig. 3. Neutralization of bath B with powdered CaO; effect of particle size.

TABLE 1

An economical balance taking into account savings and expenses related only to the mass fluxes con- nected to the treatment of I litre of spent bath

Bath A Bath B

Unit c o s t s Quantity Cost Quantity Cost (Lit/Kg) (a) (g) (Lit) (g) (Lit)

Savings: Spent bath (disposal) 210 (b) 1130 237.3 1060 222.6 HNO3 (66%) 290 (c) 62.0 18.0 31.7 9.2 Total 255.3 231.8

Expenses: CaO (powder) 112 129 14.4 104 11.6 H2SO4 (98%) 145 (c) 43 6.2 22 3.2 HF (40%) 1925 (c) 4.9 9.4 2.5 4.8 Sludges (disposal) 190 (b) 701 i 33.2 469 89.1 Steam ( 10 ata) 40 355 14.2 453 18.1 Total 177.4 126.8

Surplus budget: 77.9 105.0

(a) 1000 Lit= 0.884 US$ (Milan, 2/1/1992). (b) Transport expenses included (30 Lit/Kg). (c) Related to 1 Kg of pure substance at the stated concentration.

dur ing the neu t ra l iza t ion o f ba th B, m a d e accord ing to the chosen condi t ions . I ron and c h r o m i u m comple te ly co-precipi ta te at p H values lower than 5, while nickel is on ly slightly invo lved in this co-prec ip ta t ion (see Fig. 4) , possibly due to local p H e n h a n c e m e n t s at the surface o f the dissolving l ime particles.


25

x

. i 2o

ap

10

5

~ ~ Ni

o 2 3 4 5 6 7 9 9

pH

Fig. 4. Progressive reduction of the concentrations of metallic ions in bath B during neutralization.

If the neutralization of the bath is carried out in two stages (up to pH 5 and than up to pH 9), separated by an intermediate filtration, nickel can be recovered as Ni (OH)2, an expensive product, usable, after calcination, for the production of stainless steel. Specific tests, following this procedure, led to a nickel recovery of about 80%.

CaF2 and CaSO42H20 (as later explained, a certain SO~- content in the spent baths is a consequence of recycling the recovered acid) cannot pollute the recovered nickel hydroxide because their precipitation takes place in the first phases of neutralization together with the precipitation of Fe (OH) 3 and Cr(OH)3. A small excess of CaO in the Ni(OH)2 is not harmful because it will join the slag covering the metallic molten solution in the production of stainless steel, [ 8,9 ]. Traces of iron and chromium are tolerable as well. Fur- ther processing of the remaining sludges to recover for example iron and chromium would very likely not be economically feasible, as has already been proved in similar circumstances [ 2 ]. Without taking into account, conservatively, the recovery of nickel and using a 105-/tin sieve, the content of dry solids of the neutralized suspensions was on average 245 g/1 for bath A and 175 g/1 for bath B.

The separation of the suspended solids by sedimentation proved to be in- effective. It was therefore necessary to filter the neutralized suspension by a filter-press, capable of separating a sludge with a 55% content of liquid, [ 10 ],


a percentage largely inferior to those achievable with continuous filters. This reduces the mass of outcoming sludge as well as the amount of water necessary for its washing - taken to be twice the liquid content of the sludge.

After combining the filtered liquids with the liquids generated during washing operations, the following solutions were obtained:

A) NO~-: 50.6 g/l; Ca2+: 16.4 g/l; Fe 3+, Cr 3+, Ni 2+, F- : negligible; volume: 1228 ml per liter of spent bath.

B) NO~-: 27.2 g/l; Ca2+: 8.8 g/l; Fe 3+, Cr 3+, Ni 2+, F- : negligible; volume: 1170 ml per liter of spent bath.

It can be readily demonstrated that nearly a 15% of the nitrates, which were previously in the acid spent baths, cannot be eluted from the sludge, being linked in ferric basic nitrates, [ I 1 ], stabilized by the excess of lime, and therefore not harmful to the environment.

The concentration Through an evaluation of the overall mass balances of water and HNO3,

taking into account the discontinuous way of adding acids to the pickling tanks, and supposing a conservative 80% recovery of nitrates as HNO3 for the proposed process, the minimum allowable concentration for the recovered acid could be calculated as 10% (mass basis).

In practice, the maximum HNO3 concentration used in the pickling tanks at the ILVA plant is about 8%; the spot-addition of 10% recovered acid allows this maximum operating concentration after dilution in the whole bath to be maintained with no need of 66% (mass basis) HNO3. Spot-additions of re- generated acid in the pickling liquor, parallel to periodic draining of the spent bath could be made following the usual discontinuous adding procedures, reducing moreover the overall quantity of industrial water directly poured into the pickling solution in order to keep the iron concentration at a low level

recovered HN03 (10%) fresh HN03(66%)

RE W.Y I P,CK',NG I TREATMENT l BATH TANK J

T I 1 NO i (sludge) NO i (spent bath) drag out

N0x

Fig. 5. Mass-balance of HNO3 in case of setting up the recovery process.


water in 10% HNO 3 water in 66% HNO 3

1 F industrial water _1 PICKLING L water in 40% HF

q BATH TANK I ' ~

water in the spent b a t h evaporated water

Fig. 6. Mass-balance of water in case of setting up the recovery process.

( < 40 g/l) . Fig. 5 and 6 show the overall mass balances of nitric acid and water respectively (assumed: the whole quantity of HNO3 used is equal to 100), taking into account the studied recovery treatment.

In order to finally get a 10% HNO3 solution it is necessary to concentrate the neutralized solution up to a content of nitrates equal to 120 g/1 in view of the addition of fresh water used for washing the CaSOa2H20, which precipitates during acidification. 710 ml of water have to be boiled away from the neutralized solution deriving from 1 1 of bath A, and 905 ml from that obtained from 1 1 of bath B.

The evaporated water, once condensed, was found to contain a concentration of NO~- ions equal to 200-250 ppm, an insignificant loss for the recovery treatment but a problem for the disposability of this liquid according to Ital- ian laws. On the other hand it could be used instead of fresh water for direct addition to pickling tanks and for the washing of filtered sludges.

The acidification The main problem during acidification lies in the dissolving action of re-

covered HNO3 upon calcium sulphate, whose solubility rises up to 18.4 g/1 as HNO3 reaches the target concentration of 10% by weight, [ 12 ]. On the con- trary, due to the very low solubility of calcium fluoride, Ca 2 ÷ ions which are still part of the solution after dosing with H2SO4, are readily precipitated as HF is added in order to restore its optimal concentration, thus changing back the sulphate salt into sulphuric acid.

The small quantity of added H 2 8 0 4 does not affect the pickling efficiency of the bath except for a slight increase of its passivating power, [ 13 ], due to the presence of SOl- ions, whose aptitude for adsorption on stainless steel is well known, [ 14,15 ]. This small quantity of sulphuric acid experimentally


proved to be fairly tolerable. Fig. 7 shows the concentration of the Ca 2+ and SO4 2- ions still dissolved after a 1-hour reaction between the concentrated solution and different doses of H2SO4, and illustrates the variation of an overall cost parameter Ct defined as follows:

C, = 0.145 × (H2SO4) + 1.925 × (HF) + 0.803 × (SO 2- ),

where: (H2SO4)=dose of H2SO4 (g/liter of concentrated solution); ( H F ) = d o s e of HF (g/l c.s.; assumed equal to Ca 2+ concentration); (SO 2- ) = quantity of sulphate ions left in the solution (g/1 c.s. ).

The numerical factors are related to the unit prices of reagents and sludge disposal reported in Table 1 in accordance with the Italian market. As regards the last term, the sulphate ions present in the recovered acid solution are likely precipitated in the neutralization step of the next regenerative treatment; this implies an increase in the consumption of CaO and in the volume of outcoming sludges.

The cost function, CI, doesn't take into consideration the nature of the calcium salt preciptating during complete acidification. CaSO42H20 has a higher molecular weight than CaF2, but the latter is not easily filterable and has to be left over as a sludge with a higher percentage of humidity. Therefore the mass of the two wetted sludges is practically the same. Fig. 7 shows that the optimal dose of H2SO4 is 83 g/l, which is 87% of the stoichiometric value (95 g/l).

Continuous filters can separate the obtained gypsum with only a 40% liquid

CarSOn- Ct (g/I) 40 48 (Lit)

30

20

10

' / ? 48

0~ lJ 1.5 . ~ t r o ~ H ~ 4

Stoichiometric HpSO 4

Fig. 7. Variat ion of global cost C . of Ca 2+ and SO 2- contents in the processed l iquid after a 1 hour acidif icat ion with different doses of H2SO4.

44


content. After restoring the correct HF concentration (2.2% for bath A, 1.5% for bath B) in the acid solution derived from CaSO42H20 filration and washing, a sudden precipitation of CaF2 takes place, due to a large excess of fluoride ions. CaF2 is filterable with a 65% liquid content.

Leaving aside the excess of 40%-HF and using for the washing of calcium salt a quantity of water (twice the content of liquid in the filtered sludges) the following recovered solution was obtained: HNO3 = 106 g/l ( ~ 10% mass basis); H2504= 10.2 g/1 ( ~ 1% mass basis).

An 83% efficiency in recovering nitrates as HNO3 was on average reached so that the value 80%, assumed in Fig. 5, is to be regarded as sufficiently cautionary.

A P R E L I M I N A R Y E C O N O M I C A L B A L A N C E

On the basis of experimental data it has been possible to indicate in Table 1 an economical balance which accounts for savings and expenses, and is related only to mass fluxes. The recovery of nitric acid is compared with the disposal of spent baths (actual practice at the ILVA plant).

The item "Sludges" includes: • the neutralization precipitate, including the gypsum derived from SO42-

ions dissolved in the recovered solution. The gypsum quantity has been calculated admitting that no SO 2- ions would leave the pickling bath during pickling period. This is a cautionary assumption due to the already under- lined affinity of sulphate ions for stainless steel, which very likely makes a certain amount of them leave the pickling bath together with the stainless steel sheet;

• CaF2 and CaSOa2H20 obtained during acidification. These salts are charged with the same cost of disposal as the other sludges even if, due to their purity and chemical inertness, they could likely be disposed at a lower price.

A thermal-compressed or a multiple-stage evaporation system is set up; therefore steam consumption has been conservatively assumed as one part for every two parts of boiled-away water.

Neglecting the recovery of nickel, the bigger surplus budget evaluated for bath B (105 Lit/1 vs. 77.9 Lit / l) is mainly due to the lower sludge volume and to the price of disposal of the spent bath, fixed at the same level as more concentrated spent solutions.

Each year the ILVA plant produces 150000 tons of coils and generates 12000 m 3 of spent nitric-hydrofluoric bath. This entails an overall budget of 935- 1260 million Lit/year (814000-1097000 US$/year) left for power, labour, maintainance and depreciation costs.

150

A P R O P O S A L F O R A N I N D U S T R I A L S C A L E P L A N T

G. S A R A C C O A N D M. O N O F R I O

Fig. 8 shows a possible scheme of an industrial scale recovery plant. The equalization basin must be sufficiently large to ensure that the com-

position of the spent solution to be recovered is kept at a constant value, intermediate to those of the two tested baths.

A composition control (CC) governs the regulation of the plant (i.e. dos- age of CaO, H2SO4 and HF) .

The choice of a filter-press involves a discontinuous management of the neutralization phase, restrictable to one work shift per day (8 hours) of one worker (required expense: 60 million Li t /year= 52300 US$/year) .

The concentration and acidification phases are continuously operated, and an intermediate accumulation basin has to be set up.

The neutralization reactor can be internally lined with ebonite, a cheap an- ticorrosive gum resistant to sulphuric, nitric and hydrofluoric acids, at the determined concentrations, up to 80°C, [16]. If this temperature is ex- ceeded, an automatic cooling system is started. Highest temperatures are reached, as reported above, when the content of acids and the corrosive power of the bath has already been largely neutralized. In order to recover the liquid evaporated within the reactor, a condenser has to be set over it; NOxs passing through the condenser are piped to a NOx-reduction plant, always present in pickling plants. The reactor is automatically emptied as soon as the desired pH value is reached.

NO x abatement .,~

Plant ~ Water . . . ~ I . t.n,no -4 out,,owr_ Comp e.ed

T air

SPENT BATH STORAGE i NEUTRALIZATION i FILTRATION AND INTERMEDIATE . . . . . . . . . . . ~ . . . . . . . . j WASHING STOflA6E

Water i HF ~ -- -- HIH-~O4~ - - .. ..Steam

b. i a. !

RECOVERED HF-ACIDIFICATION H2SO4-ACIDIFICATION EVAPORATION ACID

_o

o Z -4

-in ,-4

8

C t/)

-n -4

Fig. 8. P o s s i b l e s c h e m e o f t h e r e c o v e r y p lan t .

PROCESS FOR HNOa RECOVERY FROM SPENT PICKLING SOLUTIONS 1 51

A thermal-compressed evaporator seems to be advisable for the concentration of the neutralized solution instead of a multiple-stage one, due to its flex- ibility and the relatively low flow to be treated.

For the filtration of CaSOa2H20, a continuous belt filter with counter-cur- rent washing, similar to those employed for phosphoric acid production from H 2 S O 4 treatment of phosphorite, [ 17 ], could be used. Finally, it should be verified whether CaF2 precipitation could be performed directly in the pickling tanks, as fresh acids are added, instead of setting up specific reaction and separation systems. Plant costs would be obviously lowered. The recovery of pure CaF2 is of no economic interest.

The mass of the CaF2 precipitate is rather low compared to that of the pickling sludge and presents no management difficulties.

However, very fine calcium fluoride particles could possibly be kept suspended in the softly agitated pickling solution. It should be ascertained whether or not suspended particles of CaF2 may interfere with the pickling efficiency of the bath or not. If this is a concern, the suspended CaF2 could be drained out of the pickling tanks simultaneously with the spent solution, and then allowed to settle in the storage tank together with the already mentioned pickling sludges.

The described recovery plant should not require more than 50 kW to op- erate, which entails a cost of approximately 30 million Lit/year (26100 US$/ year) for a working period of 300 days per year. 1000 million Lit/year (870000 US$/year) are thus left for maintainance and depreciation. This amount should warrant a depreciation period of no more than 3-4 years. These economical margins should be regarded as minimal due to all the conservative hypotheses assumed; they would be certainly wider if nickel were recovered in spite of a slight complication of the neutralizing section of the plant.

CONCLUSIONS

A new process for the recovery of nitric acid from nitric-hydrofluoric pickling spent solutions has been proposed. A pilot system has been tested and examined in terms of its chemical, technical, and economical characteristics with favorable results.

Its main feature lies in the concentration of a slightly alkaline liquid by the use of ordinary evaporators instead of PTFE-lined concentrators warmed by graphite heaters ([ 1,2]) or expensive membranes requiring intensive pre- treatments of the spent solution. Moreover, all kinds of nitrates (acidic or saline) could in principle be recovered (diffusion dialysis, for instance, can only recover the residual acids). On the other hand no recovery of HF is possible (F - ions are precipitated as CaF2 with no harm to the environment).

A HNO3 solution was recovered at a concentration level ( 10% mass basis)


suitable for complete recycling in the IL VA pickling process. If higher concentrations for the recovered acid were required, membrane processes, such as those based on diffusion dialysis [ 3,4 ], would probably become progressively more convenient compared to the indicated process.

In fact, transmembrane driving-forces for acid transport would be en- hanced thus reducing the membrane surface required per unit weight of recovered acid, while the concentration step in the tested process would become critical due to higher operating costs and to the risk of calcium sulfate precipitation inside the evaporator. Finally, it is worthwhile underlining that the recovery process effectively protects the environment from nitrate pollution, which was the primary goal to be reached.

REFERENCES

1 Kreppler, 1978. Recovery of Nitric and Hydrofluoric Acids from Pickling Solutions of Stainless Steel. Wire Industry, August, 639-643.

2 Anonymous, 1986. Outokumpu Continuous Annealing Furnaces and Regeneration of Pickling Acids for Annealing and Pickling Line of Stainless Steel Strips. In: Environmental High-Technology from Finland, Mexpert OY, Ministry of the Environment, Helsinki, p. 68-72.

3 Tokuyama Soda, 1988, Japanese Patent No. 63291608. 4 Sato J., Onuma M., Motomura H. and Noma Y., 1985. Recovery of Nitric Acid and Hy-

drofluoric Acid from the Pickling Solution by Dialysis. In: The Metal Finishing Society in Japan, Jitsumu Hyomen Gijutsu, Tokyo, Japan, p. 1-8.

5 McArdle J.C., Piccari J.A., and Thornburg G.G., 1989. Aquatech Systems' Pickle Liquor Recovery Process - First Year of Operation at Washington Steel, persented at the 1989 AISE Annual Convention, Pittsburgh, PA.

6 Saracco G., Onofrio M., and Genon G., 1991 Recovery of Nitric Acid from Pickling Spent Solutions, 4th World Congress Chem. Eng. Preprints; vol. 3, p. 8-19.

7 Pascal P., 1957. Noveau trait6 de chimie minerale Masson, Paris, vol. 4, p. 390. 8 Hall A.M., 1954. Nickel and Iron in Steel, J. Wiley and Sons, NY, p. 17. 9 Quenau P.E., 1961. Extractive Metallurgy of Copper, Nickel and Cobalt, Interscience, NY,

p. 287. 10 Degremont. Memento technique de l'eau, 8eme ed., Paris 1978, p. 506. 11 See ref. 7, vol. 15, p. 118. 12 Seidell A. and Linke W.F., 1958. Solubilities of Inorganic and Metal-Organic Compounds,

D. Van Nostrand Company Inc., Princeton, NJ, vol. 1, p. 668. 13 Memmi M., Berardi P.F. and Montagne G., 1976. Influenza della reattivit~ superficiale

nella operazione di bonderizzazione dell'acciaio inossidabile austenitico, Metall. Ital. 68, 6, p. 305-310.

14 Okamoto G., 1973. Passive Film of 18-8 Stainless Steel: Structure and Its Function, Corr. Sci., 13:471-489.

15 Azzeri N., 1976. Invecchiamento delle superfici passive degli acciai inossidabili, Metall. Ital., 68: 471-489.

16 Mark H.F., Bikales N.M., Overberger C.G. and Menges G. Eds., 1985. Encyclopedia of Polymer Science and Engineering, voi. 14, p. 681.

17 Pasquon I., 1980. Chimica Industriale l, CLUP, Milan, Italy, p. 270.

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A new process for HNO3 recovery from spent pickling solutions