From ESR to continuous CC-ESRR process - Gruppo Italiano Frattura

ACCIAIERIA 

From ESR to continuous CC-ESRR process: 

development in remelting technology 

towards better products and productivity 

D. Alghisi, M. Milano, L. Pazienza 

This work describes the development of the Electro Slag Remelting process at Valbruna, starting in the 

1997 with an innovative INTECO ESR plant equipped with protective gas hood (for inert atmosphere 

remelting), electrode change system and fully computer controlled. The second step was made a couple 

of years later when the plant was upgraded to ESRR ® (Electro Slag Rapid Remelting). With this new 

feature Acciaierie Valbruna was able to obtain ready to roll remelted billets (145, 160 and 200 mm 

square), getting rid of the traditional forging or blooming operations needed in case of traditional ESR 

ingot remelting. This was surely a dramatic cut off in product cost accounting and production lead time 

without loosing any of the special characteristics typical of ESR products. Unfortunately the ESRR ® 

process, very promising in terms of cycle complexity reduction and quality of the product, because 

of its “batch-type” operation, was uneconomical in regard of productivity of the plant and not feasible 

in industrial scale. The final step of this development was made at the beginning of 2002 when t 

he ESRR® plant was upgraded again and equipped with an innovative INTECO automatic manipulator, 

which resulted in a continuous process. This was the birth of the very first CC-ESRR ® 

(continuous casting electro slag rapid remelting) plant in the world. The first part of this paper focuses o 

n the development of processes and equipment, giving a brief description of ESR, ESRR ® 

and CC-ESRR® process while the second part describes the results of a series of test remelting used 

for product and CC-ESRR ® process characterization. 

Memorie 

Key words: steelshop, processes, stainless steel, solidification 

THE STANDARD ESR (ELECTRO SLAG REMELTING) PROCESS 

The standard ESR process can be summarized in the above 

picture: by remelting a consumable electrode in a superheated 

liquid slag bath a new ingot is built up in a water cooled, 

copper mould. The energy required for the melting of the 

electrode is produced by an electric current passing through 

the liquid superheated slag which is acting as an ohmic resistance. 

The steel, which melts off from the electrode tip drops 

through the hot reactive liquid slag thus forming ideal 

conditions for slag metal reactions. 

The constant metal level in the mould is realized by means 

Fig 1 – ESR Process 

schematics. 

Fig. 1 – Schema del processo 

ESR. 

Davide Alghisi, Michele Milano, 

Laura Pazienza 

Acciaierie Valbruna S.p.a. - Vicenza 

Paper presented at the 2 nd International 

Conference on New Developments 

in Metallurgical Process Technology, 

Riva del Garda, 19-21 September 2005, 

organised by AIM 

1/2005 

la metallurgia italiana 21


1/2005 Memorie 

of a retracting baseplate, whose downward movement is set 

by the control system according to the meltrate and therefore 

ingot growth. 

The ingot length is limited by the maximum baseplate range 

of 4 meters; as soon as the baseplate reaches the bottom position, 

the remelting process is stopped to remove the ingot. 

The maximum capacity of the plant is limited to a 7 ton ingot 

produced out of five electrodes in sequence in a 520 mm 

round mold. 

The 520 Ø mm ingots need to be forged or bloomed in a 

roughing mill to be ready for hot rolling long products as the 

ones produced by Valbruna. 

Unfortunately the forging or blooming operation, in addition 

to be a further cost in terms of production cost accounting 

Fig 2 – The complete ESR 

production cycle. 

Fig. 2 – Il ciclo completo della 

rifusione ESR. 

Fig. 3 – ESRR ® process 

schematics. 


ESRR ® . 

and lead time, eliminates some of the structural characteristics 

of the remelted product, for example the axial growing 

direction and the grain dimension. 

THE ESRR (ELECTRO SLAG RAPID REMELTING) PROCESS 

The aim of the ESRR® process is to remelt near-net shape 

billets that can be directly hot rolled without any additional 

forging or blooming operation. 

This leads to a dramatic economic advantage and a real lead 

time reduction due to the elimination of some process phases: 

• Reheating + Forging + grinding 

• Reheating + Blooming + grinding. 

22 

la metallurgia italiana


Memorie 

Fig. 4 – The ESRR ® mould during process. 

Fig. 4 – Schema della lingottiera ESRR durante il processo di rifusione. 

Fig. 5 – The complete mould assembly during CC-ESRR ® process. 

Fig. 5 – La lingottiera ESRR. 

Fig. 6 – Cobalto 60 scintillator. In the photo the device is covered by 

a yellow chassis for radiation protection. 

Fig. 6 – Lo scintillatore al cobalto 60. Nella foto il dispositivo è 

coperto da un carter di sicurezza. 

Fig 7 – Details from the new 145mm ESRR T-shaped mould body. 

Details of the particular distribution system of the cooling jacket. 

Fig. 7 – Dettagli del corpo lingottiera da 145 mm per il processo 

ESRR. Dettaglio del particolare sistema di raffreddamento a 

camicia. 

Furthermore the “as cast” remelted billet shows a structural 

direction of solidification and a fine and uniform grain size 

that are ideal for hot rolling. Those particular features were 

(as seen before) generally lost after reheating and forging 

operations. 

The core of the electro slag rapid remelting process is indeed 

the particular copper T-shaped square mould whose design 

has been optimised (by using a FEM analysis) to 

perform a proper heat transfer. The lower-narrow and upperwider 

part of the copper mould are both water cooled to 

perform a constant and uniform heat subtraction over the 

whole mould. 

During ESRR ® process, by remelting a consumable electrode 

in a superheated slag bath (1), a new ingot is built up in a 

water cooled copper mould. The energy required for melting 

the electrode is produced by the electric current passing through 

the liquid superheated slag which is acting as an ohmic 

resistor. 

The liquid metal droplets from the electrode tip are collected 

(2) in the narrow, lower part of the mould, where the initial 

solidification takes place and the remelted billet (3) is 

continuously formed. 

The remelted billet is withdrawn (as in standard ESR operation) 

by means of a retractable base plate. 

The baseplate movement can’t be controlled only by the 

standard meltrate signal (as in standard ESR), as the position 

of the metal level must be safely kept under the T-shaped 

extension. 

In fact, if the metal would solidify in the wider part of the 

mould it would be impossible to withdraw the billet, the 

process would have to be interrupted and the whole mould 

assembly has to be dismantled. 

Therefore a new and more precise and reactive signal was 

installed to control the baseplate retraction. 

The signal for controlling the liquid metal level is generated 

by a radioactive Co-60 isotope scintillator. 

As shown on the process scheme both measuring devices 

are installed at the top end of the narrow part of the 

mould. If the metal level (2) grows too high, the signal, 

which is sent by the radioactive source and detected by a 

scintillator installed at the opposite side of the mould, will 

be reduced. 

Such a signal change forces the PCS to retract the base 

plate resulting in a lowering of the metal level. 

As in the standard ESR, the baseplate retraction is limited 

by the pit depth, resulting in a maximum billet length of approx. 

4 meters. The net length will be a little bit shorter due 

to losses at the bottom and top of the billet. 

Furthermore it is impossible to remove the billet without 

breaking the power circuit, so only one billet can be produced 

at a time out of one single electrode. 

The ESRR ® process turned out to be a fundamental, primary 

testing phase, a first step for further development. This phase 

was very important for three main reasons: 

• Plant-design optimization: For better understanding of the 

heat transfer mechanism and to receive further ideas regarding 

optimization of the mould design to obtain better 

quality billets without deformation or internal stress, a 

FEM thermal analysis has been made in collaboration 

with AVL LIST. 

• Process design optimization: An innovative deslagging 

device has been developed. Various attempts were made 

in the past to remove the slag from the mould as long as it 

is still liquid. The latest utilized method consists in a slag 

1/2005 




Fig. 9 – FEM analysis of the mould during the process: the mould 

temp. distribution. Photo Courtesy of INTECO / AVL LIST. 

Fig. 9 – Analisi agli elementi finiti della lingottiera durante il 

processo. 

vacuum suction device operating from the mould top 

before the removal of the billet 

• Remelted billets property definition and comparison to the 

traditional ESR product: the excellent results have been 

described in another work, in any case the quality of the 

billets is comparable (and in many cases better) to the traditional 

ESR-ingots in terms of soundness, inclusion contents 

as well as physical and chemical properties. 

Out of the experience gained in the batch-type ESRR ® -operation 

the next step, the CC-ESRR ® process has been developed! 

Fig. 8 – The complete ESRR ® 

production cycle. 

Fig. 8 – Il ciclo completo della 

rifusione ESRR ® . 

Fig. 10 – The slag suction device ready to operate. 

Fig. 10 – Il sistema di aspirazione della scoria. 

THE CC-ESRR® (CONTINUOS CASTING ELECTRO SLAG 

RAPID REMELTING) 

In the beginning of 2002, after a plant revamping, INTECO 

has equipped the ESRR ® plant with a special automatic manipulator, 

which results in a continuous Rapid Remelting 

process (CC-ESRR ® ). 

In the CC-ESRR ® the baseplate for the retraction of the billet 

is replaced by 2 drive and 4 guiding rolls, whose movement 

is again managed by the Control System based on the 

signals of the metal level detector. 

24 



Fig. 11 – CC-ESRR ® process 

schematics. 


CC-ESRR ® . 

Memorie 

Fig. 12 – The complete 

CC-ESRR ® production cycle. 

Fig. 12 – Il ciclo completo 

della rifusione CC-ESRR ® . 

As soon as the billet reaches its required length, a powder cutting 

torch (oxygen + iron powder injection) automatically 

cuts the billet which is then removed in an unloading position. 

The revolutionary innovation is that all the continuous production 

of up to 40 billets in one remelting sequence out of 

multiple of electrodes. 

This results in a decrease of the slag, tooling and starting costs 

due to the continuous production by an increase of the 

productivity of the plant. 

This is surely the definitive step versus a price competitive 

product that incorporate all the advantages of ESRR® product 

with the advantages (in terms of volumes and economic 

impact) of continuous production. 

The excellent results in terms of process efficiency makes 

the CC-ESRR® the definitive development of the ESR process 

for producing hot rolled long products. 

The product remelted with the continuous process incorporates 

in fact all the advantages of the ESR process (in regard 

of quality of the metal), of the ESRR® process (in regard of 

reduced cycle complexity and overall process lead-time) 

and the efficiency, volumes and low-running-cost of the 

continuous processes. 

The core of the process is once more the T-shaped mould 

that is exactly the same that is used during batch-type 

ESRR ® process. 

The automatic withdrawal manipulator 

The main innovation in the CC-ESRR® plant is indeed the 

automatic manipulator. The manipulator must: 

• Retract and guide the produced billet 

• Keep the electrical contact to close the power circuit 

• Cut the billet as soon as it reaches the required length. 

1/2005 




Fig. 13 – The CC-ESRR ® manipulator schematics. 

Fig. 13 – Schema del manipolatore CC-ESRR ® . 

Fig. 14 – The CC-ESRR® manipulator during process: 

1. Brushless electrical engines 

2. Electrical contacts and driven rolls 

3. Guiding rolls 

4. Cutting device chamber 

5. Retracting baseplate with billet holder 

6. Manipulator body mounted on a rail system. 

7. Operating platform. 

Fig. 14 – Il manipolatore durante il processo: 

1. Motori elettrici brushless 

2. Contatti eletttici e rulli motrici 

3. Rulli guida 

4. Camera del dispositivo di taglio 

5. Piastra base retrattile con alloggiamento della billetta 

6. Il manipolatore in posizione di lavoro 

7. Piattaforma di lavoro. 

For an easy and quick change between standard ESR- and 

CC-ESRR ® -operation the manipulator is installed on a rail 

system to move between an operating position (CC-ESRR ® ) 

and a parking position (standard ESR). 

The driving-guiding rolls 

To drive and guide the billet out of the mold, the manipulator 

is equipped with three pairs of rolls. One pair is driven 

and its movement is controlled by the PCS system according 

Fig. 15 – The complete sliding contact assembly. 

Fig. 15 – Il gruppo dei contatti striscianti. 

Fig. 16 – The brushless electrical engines that drive the rolls. 

Fig. 16 – I motori brushless che azionano i rulli. 

Fig. 17 – Overview of the cutting device, with the swivelling torch, 

the water spray system for the granulation of the generated slag 

and protection of the copper box. 

Fig. 17 – Vista del dispositivo di taglio con la torcia , il sistema di 

granulazione delle polveri e la protezione di rame. 

to the metal level signal. The other two pairs are used only 

to guide the billet out of the mold. 

Furthermore, the driven rolls are equipped with a set of innovative 

sliding contacts whose task is to close the power 

circuit formed by the electrode, the slag, the billet and the 

bus bar system. 

The cutting device 

The cutting device is an oxygen torch equipped with an iron 

26 



Fig. 18 – The cutting device during operation. Note the billet and 

the billet holder are mounted on the retracting baseplate. 

Fig. 18 – Il dispositivo di taglio durante il processo: notare la 

billetta e il porta-billetta montato sulla piastra base retrattile. 

Fig. 19 – A sample slice cut by the oxygen-powder torch. Note the 

good surface. 

Fig. 19 – Esempio dell'ottima superficie si una fetta di billetta 

dopo il taglio. 

powder injection system, and is capable of cutting a 145 mm 

billet in less than a minute. 

During the cutting operation, all the fumes and the slag generated 

by the process are collected within a fume hood with 

a special water spraying system. 

As said before, in this case the billet holder is mounted to 

the retracting baseplate and moves the remelted billets (immediately 

after cutting) into the unloading position. 

• Influence of a grade change during CC-ESRR ® operation: 

it was evaluated, how the interface between two steelgrades 

looks like in case of a grade change during one remelting 

sequence 

• Product characterization of the as cast billets: including 

chemical, metallographical and physical properties evaluation 

• Product characterization of the hot rolled billets: including 

chemical, metallographical and physical properties evaluation. 

GRADE CHANGE BETWEEN AISI 304 AND AISI 660 

The change of electrode from AISI 304 to AISI 660 during 

one melting sequence was evaluated. 

The test (with the aim to evaluate when the change of grade 

is completed) has been carried out on the following grades 

and process parameters. 

Grade Melting rate Step Back step Slag 

(Kg/h) (mm) (mm) 

AISI 304 500 4 0 CAF3 + 15 % Ti 

AISI 660 460 4 0 CAF3 + 15 % Ti 

Table 1 –Test process parameters. 

Tabella 1 – Parametri del processo di test. 

The billet has been cut to verify the transition between the 

two grades. 

As the grade change is combined with an electrode change 

and therefore a power interruption the transition area and 

shape of the pool depth was clearly visible. It amounted to 

approx. ??? mm. 

This part has to be cut out to ensure no mixing of the two 

grades at the top or bottom of the respective billets of different 

grades. 

AISI 403 CC-ESRR® BILLET ANALYSIS 

Four billets of AISI 403 have been examined as representative 

of 145x145 mm production. 

Memorie 

OPERATIONAL TESTS RESULTS 

A lot of tests have been executed in the CC-ESRR ® plant, in 

order to define process parameters as well as product characteristics. 

The grades used in those tests are commonly 

used and appreciated by our customers. Therefore a direct 

comparison of the new CC-ESRR ® results to already achieved 

values (ESRR ® and ESR) in terms of product quality 

and process parameters was possible. 

• AISI 403: A martensitic steel grade used mainly for steam 

turbine blades. 

• AISI 304L: An austentic steel grade used for corrosion resistant 

component 

• WN 1,4980: Nickel alloy grade used for high temperature 

corrosion resistant bolts (Ti alloyed). 

The following evaluations have been executed: 

Fig. 20 – Transversal and longitudinal section of the transition 

zone. 

Fig. 20 – Sezione trasversale e longitudinale della zona di 

transizione. 

1/2005 




Table 2 – Process parameters, during ESRR ® operation. 


AISI 403 FLAT BARS 

FROM CC-ESRR® BILLET PRODUCT ANALYSIS 

One batch of AISI 403 has been manufactured employing 

billets coming out of the first CC-ESRR ® campaign. 

The final products are flat bars (117.47 x 41.27 mm) for 

steam turbine blades. 

Melting rate Step Back step 


Slag 

500 6 0 CAF3 

Table 3 – Process parameters, during ESRR ® processing. 


The results of these first tests show a product that fully 

meets our customer requirements. During the standard processing, 

no anomalies have been observed to our standard. 

Surface anomalies observed during visual examination occurred 

during our manufacturing and not for CC-ESRR ® 

process. 

• Mechanical properties: Mechanical properties meet the 

required characteristics. 

• Metallographic examinations: Metallographic examination 

after hardened + tempered shows a tempered structure 

with a grain size around 6 –7 ASTM and a percentage 

of d ferrite of about 0.5 %. 

• Impurity level: The level of impurities observed is very 

low. According to E45, the worst inclusion size was 1.5. 

according to UNI 3288-80, K1 is 0. 

• Macrographic examinations: Macrographic examinations 

show a fine grain, no segregations or other anomalies have 

been observed. 

Chemical analysis 

(See Tab. 4) 

Source melt Cast Product 

analysis analysis analysis 

C 0.135 0.128 0.13 

Si 0.33 0.30 0.28 

Mn 0.44 0.43 0.42 

Cr 12.05 12.17 12.12 

Ni 0.45 0.44 0.44 

Mo 0.12 0.11 0.11 

Cu 0.11 0.08 0.08 

Al 0.009 0.035 0.032 

V 0.072 0.07 0.07 

Co 0.022 0.02 0.02 

P 0.017 0.017 0.018 

S 0.0010 0.001 0.001 

Sn 0.0060 0.005 0.004 

N 0.0350 0.0410 0.0410 

Table 4 – Chemical analysis. 

Tabella 4 – Analisi chimica. 

Impurity level 

Evaluation according to: UNI 3244-82 K1 = 0 

Evaluation according to: ASTM E45 

A B C D 

T H T H T H T H 

Required 2.0 2.0 2.0 2.0 2.0 2.0 2.0 2.0 

Found 1.0 1.5 

Table 5 – Impurity level according to ASTM E45. 

Tabella 5 – Inclusioni secondo l'ASTM E45. 

Mechanical characteristics 

(See Tab. 6) 

Hardening at 980°C → air 

Tempering at 700°C → air 

AISI 304 CC-ESRR® BILLET ANALYSIS 

Two batches of AISI 304 have been manufactured employing 

two billets coming from the first CC-ESRR® campaign. 

The final products are square bars (40x40 and 70x70). 

Melting rate Step Backstep 


Slag 

500 6 0 CAF3 

Table 7 – Working parameters, during ESRR processing. 


The results of these first tests show a product that fully 

meets our customer requirements. 

During the standard processing, no anomalies have been 

found out from our standard. Visual inspection after hot rolling 

shows some marks. These marks, after pickling, could 

not be found out. 

• Mechanical properties. Mechanical properties meet the required 

characteristics. 

• Metallographic examinations 

- Hot rolled conditions: grain size is correct. 

28 



Table 6 – Mechanical characteristics. 

HB 

Rm Rp 0.2 E% RA % KV 

(N/mm 2 ) (N/mm 2 ) 

Required according to Customer 1 240 max 850 max 560 ÷ 660 15 min 50 min 35 min 

Required according to Customer 2 217 – 248 690 min 550 min 20 min 60 min 81 min 

Hard & temp in lab furnace 243 770 615 25.3 68.6 134 137 143 

Hard & temp in manufacturing 230 729 569 25 67 126 131 138 

Tabella 6 – Caratteristiche meccaniche. 

Memorie 

Fig. 21 – Micrographic 

examinations. 

Fig. 21 – Esami micrografici. 

- Solution treated in laboratory furnace: grain size is correct. 

- Solution treated in manufacturing furnace: It can be remarked 

that after solution treatment, this material recrystallize 

with a grain finer than in the hot rolling condition. 

Grain size is correct. 

• Macrographic examinations. Macrographic examinations 

show a fine grain, no segregations or other anomalies have 

been found out. 

Chemical analysis 

(See Tab. 8) 

Mechanical characteristics on 40x40 square 

(See Tab. 9) 

Mechanical characteristics on 70x70 square 

(See Tab. 10) 

SPECIAL THANKS 

Many thanks to everybody at Valbruna’s laboratory, quality 

department and ESR plant for their time and concrete help 

Source melt Cast Product 

analysis analysis analysis 

C 0.036 0.040 0.036 

Si 0.53 0.46 0.46 

Mn 1.48 1.46 1.46 

Cr 18.16 18.17 18.17 

Ni 8.07 8.11 8.11 

Mo 0.44 0.44 0.44 

Cu 0.43 0.44 0.44 

Al 0.012 0.050 0.050 

Co 0.10 0.10 0.10 

B 0.0023 0.0017 0.0017 

P 0.027 0.027 0.027 

S 0.0010 0.0010 0.0010 

Sn 0.0110 0.010 0.010 

N 0.0672 0.0755 

Table 8 – Chemical analysis. 

Tabella 8 – Analisi chimica. 

1/2005 





examinations on 40x40 square. 

Fig. 22 – Esame micrografico 

su un quadrato 40x40. 

HB Rm (N/mm 2 ) Rp 0.2 

(N/mm 2 ) E% RA % 

Required 515 min 205 min 30 min 40 min 

Solution treated + cold drawn 240 689 485 49.4 75.4 

Solution treated in lab at 1080°C → wq 600 272 66.3 78.5 

Table 9 – Mechanical 


Tabella 9 – Caratteristiche 

meccaniche. 

HB core HB _R HB surface 

Hot rolled condition 230 230 240 

solution treated in laboratory at 1080°C → wq 180 200 192 

HB Rm (N/mm 2 ) Rp 0.2 

(N/mm 2 ) E% RA % 

Required 215 max 500 - 700 190 min 45 min 

Solution treated in manufacturing 175 590 288 64.0 78.8 

Solution treated in laboratory at 1080°C → wq 598 272 65.2 77 

Table 10 – Mechanical 


Tabella 10 – Caratteristiche 

meccaniche. 

HB core HB _R HB surface 

Hot rolled condition 180 187 195 

solution treated in laboratory at 1080°C → wq 190 194 176 - 162 

30 




examinations on 70x70 square. 

Fig. 23 – Esame micrografico 

su un quadrato 70x70. 

Memorie 

for the making of this work. Special thanks also to Monika 

Boh and Harald Holzgruber at INTECO for their great help 

and useful advices. 

REFERENCES 

• AA.VV. “ESR Know How”, INTECO Technical Manual, 

INTECO, Bruck a.d.Mur, 1997 

• D.Alghisi, “ The ESR process under protective atmosphere”, 

“METEC 99 Inteco Symposium proceedings”, Düsseldorf, 

June 1999. 

• D.Alghisi, “ The ESRR process under protective atmosphere”, 

“Medovar Memorial Symposium 2001 proceedings”, 

Kyiv, March 1999. 

• H.Holzgruber, “The electroslag rapid remelting”, “ME- 

TEC 99 Inteco Symposium proceedings”, Dusseldorf, June 

1999. 

• Ennemoser, H.Petrin, “ESR mould CFD-FEM thermal 

analysis”, AVL Report, September 1998. 

• W.Holzgruber, “Production of high-quality billets with 

ESRR process”, INTECO, Bruck a.d.Mur, 1997 

• Photographic reference: 

M. Boh, INTECO G.m.b.H; 

D. Alghisi, Acciaierie Valbruna S.p.a. ESR archive. 

AVL List Report. 

Fig. 24 – The very first billets produced sequentially by CC-ESRR ® . 

Fig. 24 – Le prime billette prodotte col sistema CC-ESRR ® . 

1/2005 




FROM ESR TO CONTINUOUS CC-ESRR PROCESS: 

DEVELOPMENT IN REMELTING TECHNOLOGY TOWARDS BETTER 

PRODUCTS AND PRODUCTIVITY 

PAROLE CHIAVE: acciaieria, processi, acciaio inox, 

solidificazione 

Questo lavoro descrive l'evoluzione del processo ESR presso 

le Acciaierie Valbruna SpA a partire dell'installazione 

dell'innovativo impianto ESR (dotato di campana protettiva 

per la rifusione in atmosfera inerte) Inteco del 1997. 

Il secondo passo è stato fatto un paio d'anni dopo con l'implementazione 

nell'impianto del processo ESRR® (Electro 

Slag Rapid Remelting). Con questa nuova caratteristica 

l'impianto è stato messo in grado di produrre billette (da 

145, 160, 220 mm) pronte per la laminazione saltando a piè 

pari i processi di fucinatura o blumatura del lingotto rifuso 

necessari nel processo ESR tradizionale. 

Questo processo è stato senza dubbio un incredibile passo 

avanti in termini di diminuzione della complessità dei cicli e 

ABSTRACT 

del lead time senza per questo perdere nessuna delle caratteristiche 

qualitative tipiche del prodotto ESR. Sfortunatamente 

per quanto innovativo e promettente dal punto di vista 

tecnico, il processo ESRR non poteva essere che una 

"tappa intemedia" in un cammino di industrializzazione del 

processo: infatti a causa del suo approccio tipicamente 

"batch" era in grado di produrre una billetta alla volta con 

effetti negativi sulla produttività e sul costo di processo. 

Il passo finale di questa evoluzione è stato fatto alla fine del 

2002 quando l'impianto è stato equipaggiato con un innovativo 

manipolatore automatico in grado di rendere il processo 

ESRR continuo. Era nato il primo impianto su scala industriale 

di CC-ESRR (continuous casting electro slag remelting) 

al mondo 

La prima parte di questo lavoro è focalizzata sulla descrizione 

dello sviluppo del processo e dell'impianto con attenzione 

particolare alle tappe innovative tra i vari stadi del 

processo (ESR-ESRR-CCESRR), mentre invece la seconda 

parte prende in esame i risultati sperimentali su alcuni prodotti 

utilizzati per la caratterizzazione del processo. 

32 


PROCESSI 

Modern secondary cooling technology 

in continuous casting of steel 

R. Boyle, J. Frick 

Continuous casting machines are now required to cast a wide range of steel grades while maximising 

production output. Consistent production of prime quality product requires increased operational and 

maintenance flexibility of the caster so that the optimum casting parameters can be maintained for each 

steel grade. This flexibility extends not only to the machine elements and control systems, but also to the 

secondary cooling system and demands more efficient and reliable spray cooling. Attentive design of 

secondary cooling systems, through cooling zone positioning, nozzle layout, nozzle selection and the use 

of appropriate secondary cooling control systems can provide these requirements. Minimum down time 

for maintenance is a key factor in maximising caster production; the use of the latest piping header 

systems and nozzle mounting arrangements can contribute to minimum secondary cooling maintenance. 

These header systems provide a rigid and self-aligning mounting for the nozzle, ensuring both nozzle 

alignment and maintenance accessibility. 

Memorie 

Key words: secondary cooling, air mist nozzle , water distribution, vertical piping design 

INTRODUCTION 

Continuous casting machines are now required to cast a wide 

range of steel grades, in particular slab casters must cast 

steels ranging from ULC and low carbon grades to high carbon 

and high quality pipeline grades. This must be achieved 

while maximising production output. Consistent production 

of prime quality product requires increased operational and 

maintenance flexibility of the caster so that the optimum casting 

parameters can be maintained for each steel grade. 

This flexibility extends not only to the machine elements 

and control systems, but also to the secondary cooling system 

and demands more efficient and reliable spray cooling. 

Of particular concern when designing a secondary cooling 

system are: 

• Steel grades to be cast and their casting speeds. 

• The roll support geometry and machine segment layout. 

• Ease of maintenance. 

• Secondary cooling control systems. 

This paper focuses on the design of a secondary cooling system 

that uses the latest nozzle technology to fulfil the production 

requirements of today’s casters. Unlike in the early 

days, the layout of the secondary spray cooling system is 

one of the first steps when a new continuous casting machine 

is designed, or when an existing machine undergoes a 

major revamp. 

NOZZLE LAYOUT 

Ray Boyle - Lechler Ltd., UK 

Juergen Frick - Lechler GmbH, Germany 

Paper presented at the 2 nd New Developments in Metallurgical Process Technology, 

Riva del Garda, 19-21 September 2004, organised by AIM 

Fig. 1 – Staggered 

nozzle arrangement. 

Fig. 1 – 

Disposizione 

sfalsata degli ugelli. 

Fig. 2 – Spray width control – alternating nozzles. 

Fig. 2 – Controllo della larghezza dello spray – ugelli alternati. 

A good nozzle layout is paramount in fulfilling the operational 

and production requirements. It is essential that nozzle 

arrangements produce an even heat removal across the 

strand while maintaining a stable spray pattern. Spray collision 

with support rolls should be avoided as this will result 

in inefficient use of spray water and a reduction in heat transfer. 

Generally multi-nozzle layouts should be the preferred 

arrangement. In the final area of solidification of non critical 

steel grades, typically the horizontal section of curved casters, 

it is possible to reduce the number of nozzles in a roll 

gap to one or two as this is a less critical area for solidification. 

The staggering of nozzle pairs in consecutive roll gaps, 

see Figure 1 will ensure even surface temperatures. 

Spray width control can be achieved with a multi-nozzle 

configuration. In a multi-nozzle arrangement the outermost 

nozzles are systematically turned off in relation to the strand 

width as shown in Figure 2 where a nozzle layout which alternates 

the number of nozzles in consecutive roll gaps can 

be used. If a more finer control is required then an inline arrangement 

as shown in Figure 3 can be used. 

Heat removal from the strand is not only a function of spray 

cooling; other mechanisms are also prevalent, for example 

heat removal by the support rolls. Heat removed by rolls can 

have a significant effect on the strand surface temperature 

1/2005 


PROCESSI 


Fig. 3 – Spray width control-inline nozzles. 

Fig. 3 – Controllo della larghezza dello spray – ugelli allineati. 

Fig. 4 – Even heat removal by rolls and sprays. 

Fig. 4 – Eliminazione omogenea del calore mediante rulli e spray. 

and strand solidification conditions. If the heat removed by 

rolls is considered even across the strand width together 

with even heat removal by the sprays then ideal solidification 

conditions as shown in Figure 4 should exist. 

NOZZLE SELECTION 

Nozzle selection can only occur after derivation of the solidification 

profiles, the cooling zone layout and the nozzle 

layout. The heat extraction required to achieve the solidification 

profiles is converted into cooling zone water flows 

using nozzle heat transfer coefficients. The specific nozzle 

flows can then be derived from the maximum water flow associated 

with each cooling zone. Prior to final nozzle selection, 

operational factors must be considered, these factors 

include: 

• Spray water temperature compensation factor – heat extraction 

capability reduces with increasing water temperature, 

see Figure 5 and can lead to both operational and 

quality problems. Typically this factor would be applied 

during hot summer periods. 

• Increased flow rate on the outer radius – this is used to 

equalise cooling on both the inner and outer strand faces 

by compensating for the gravitational effect on the outer 

radius. 

• Safety factors – any allowance above the calculated water 

flows. 

It is important for today’s caster designers to have access to 

nozzles which have high turndown (control range, min./ 

max. water flow) capabilities for not only operational reasons 

but also to minimise the nozzle varieties in one particular 

machine. Both maintenance and inventory managers ap- 

Fig. 5 – Effect of spray water temperature on HTC. 

Fig. 5 – Effetto della temperatura dell’acqua dello spray. 

Fig. 6 – Pressure flow diagram of Mastercooler air mist nozzles. 

Fig. 6 – Diagramma di flusso degli ugelli per la nebulizzazione 

d’aria Mastercooler. 

preciate this effort. Latest achievements in air mist nozzle 

research and development are now providing designs with 

turn down ratios wider than ever before and with lower air 

consumption. The flow diagram in Figure 6 of a Lechler 

Mastercooler air mist nozzle shows that a turn down ratio of 

more than 1:20 at a constant air pressure of 2.5(bar) is not 

impossible between water pressures of 0.5(bar) and 7(bar). 

NOZZLE DESIGN 

Within the last 5 years the vertical segment piping developed 

for the Lechler Mastercooler air mist nozzles with vertical 

square header pipes has almost become an industry standard 

design. The air mist nozzles now equipped with plates are bolted 

vertically onto adapter plates as shown in Figure 9. Once 

the secondary cooling system layout is completed and the mechanical 

design of the segments is known, it is the spray nozzle 

manufactures task to design nozzles which provide a 

uniform water distribution across the strand surface and over 

the entire turn down ratio. Tolerances of ± 15% from the mean 

value can be achieved with a multi nozzle arrangement at water 

pressures between 1.0 and 7.0 bar. Figure 7 shows the water 

distribution measurements of a twin nozzle arrangement. 

The nozzle pitch is 400(mm), the spray height 200(mm), and 

the air pressure 2(bar) constant and 7(bar) water pressure. All 

nozzles are mounted outside of the framework at the rear side 

of the segment with only the nozzle pipe, carrying the spray 

tip, extending down to the spray position. A very rigid header 

pipe giving a nozzle self-alignment is the result; the nozzle 

spray position is always secured. A “Hose less” fluid supply 

system becomes a standard. Frequent replacement of many expensive 

water and air hoses is no longer required. 

50 


PROCESSI 

Fig. 7 – Uniform water distribution measurement. 

Fig. 7 – Misura della distribuzione uniforme dell’acqua. 

Fig. 9 – Mastercooler mounted on vertical plate and square pipe 

manifold. 

Fig. 9 – Mastercooler montato su collettore con piastra verticale e 

square pipe. 

Memorie 

Fig. 8 – Header manifold old design with many small feed pipes. 

Fig. 8 – Vecchio tipo di collettore con molti piccoli tubi di 

alimentazione. 

MAINTENANCE 

Because of their internal mixture, air mist nozzles require 

two separate feed pipes for compressed air and water. Until 

recently small diameter pipes were used to feed both the 

fluids and to hold the nozzles in place. Only in special cases, 

where one fluid was fed directly by a hose, additional supports 

were provided. Nozzle staggering between the roller 

gaps within one segment becomes much easier since different 

nozzle positions can be served from only one header pipe 

manifold, see Figure 10. The conventional air mist nozzles 

mounted on these small pipes are hidden inside the segment 

framework as shown in Figure 8. 

Having the nozzles mounted so close to the strand makes 

maintenance (cleaning or adjustment) impossible unless the 

segment is removed from the machine. In the case of a break 

out nozzles must be completely replaced, which is very costly. 

Strand surface defects can often be traced back to misaligned 

spray nozzles. Header pipes such as shown are one 

source of such misalignments. The many small air and water 

pipes are often out of position due to mechanical impact or 

thermal reasons. The large number of small, individually 

bent pipes, are also expensive to manufacture. The nozzles 

and header pipes with the vertical plate connection are also 

an ideal solution for beam blank casters with air mist cooling. 

Instead of a complex manifold only two square header 

pipes with two nozzles bolted on, are required. The advantages 

described above are also true here. The bends of the 

nozzle extension pipes can be made to suit. With the aid of 

the “Split pipe” design the two nozzles on either side (Pos. 1 

and 2 of Figure 11) can be identical with the front pipe turned 

by 180°, hence one nozzle type can serve both positions 

in one gap. 

Fig. 10 – Staggered Mastercooler nozzles served from one pipe 

manifold. 

Fig. 10 – Ugelli Mastercooler con disposizione sfalsata serviti da 

un collettore a tubo singolo. 

Fig. 11 – Mastercoolers installed in a beam blank caster. 

Fig. 11 – Mastercoolers installato in un impianto di colata di 

bramme. 

1/2005 


PROCESSI 


Fig. 12 – The new Lechler Billetcooler full cone air mist nozzle. 

Fig. 12 – Il nuovo ugello nebulizzatore conico Billetcooler 

Lechler. 

NEW AIR MIST NOZZLES FOR BILLET AND BLOOM CASTING 

When air mist cooling becomes necessary for a billet or 

bloom caster, flat jet nozzles may not always be the best 

choice. This is especially true where the formation of 

Halfway Cracks may be experienced. This type of crack has 

been shown to be caused by reheating of the strand surface 

after it has passed the sharp heat extraction zone beneath a 

spray jet. During this reheating process the surface expands 

and imposes a tensile strain on the hotter and weaker inner 

material, which can then crack. The use of flat jet nozzles 

intensifies this effect. Full cone nozzles or oval cones provide 

a softer cooling by extracting heat over an extended surface 

area. These two spray patterns are the standard for single 

fluid water secondary cooling systems, however there 

has not been an adequate version using air mist. Common 

full cone air mist nozzles show unstable spray performances, 

very high air consumptions and a tendency to clog very 

easily. Oval cone air mist nozzles are often flat jet nozzles 

with multi slot orifices. Non uniform spray patterns and the 

TECNOLOGIA MODERNA DI RAFFREDDAMENTO SECONDARIO 

NELLA COLATA CONTINUA DELL’ACCIAO 

PAROLE CHIAVE: 

processi, acciaio, colata continua 

ABSTRACT 

very narrow easy to clog slots, made these nozzles barely 

more than a compromise. With the new Lechler Billetcooler 

Figure 12, a new generation of full and oval cone air mist 

nozzles it is now possible to utilise air mist cooling in billet 

and bloom casters as effectively as in slab casters. The compact 

block design allows mounting both on horizontal spray 

bars and on vertical “Banana” nozzle headers, Turn down 

ratios as wide as 1:14 have been achieved at water pressures 

between 1.0 and 10.0(bar) at 2(bar) air constant. Nominal 

spray angles for circular full cone nozzles range between 

60° and 90°. Free passages with 2.0mm in diameter are approximately 

three times larger than before for a nozzle size 

with flows ranging from 0,5(l/min) at 1(bar) water pressure 

and 5.0(l/min) at 7(bar) water pressure at a constant 2(bar) 

air pressure. Extremely critical cooling problems have successfully 

been solved in a 5-strand bloom machine casting 

more than 250 steel grades and also critical stainless steel 

rounds. 

CONCLUSION 

The benefits for the user but also for the machine designer 

described in this paper are well established facts. The most 

important of them are: o Reduced incidence of surface defects 

and crack formation o Reduced maintenance and operation 

costs o Improvement of operation safety o Enlargement 

of caster product mix o Increased caster production 

The modern air mist nozzle and header pipe technology can 

be incorporated into new machines as well as into existing 

casters for billets, blooms, beam blanks, slabs and thin slabs. 

REFERENCES 

1) Jürgen Frick. “USER BENEFITS OF MODERN AIR 

MIST NOZZLE AND SECONDARY COOLING SY- 

STEM TECHNOLOGY” 85th Steelmaking Conference 

March 10-13, 2002 Nashville, Tennessee USA 

2) Jürgen Frick, Roman Haap “Improved Secondary Cooling 

in Continuous Casting” at XXXII Seminario de Fusao, 

Refino e Solidificacao dos Metais on May7th & 8th , 

2001 in Salvador / Brazil 

Gli attuali impianti di colata continua devono essere in grado 

di fondere un’ampia gamma di tenori di acciaio, massimizzando 

nel contempo la capacità produttiva. Una produzione 

efficiente di prodotti di prima qualità necessita una 

maggiore flessibilità operativa e una maggiore manutenzione 

in modo da mantenere parametri di colata ottimali per 

tutti i tipi di acciaio. Questa flessibilità non riguarda solo 

agli elementi dell’impianto e ai sistemi di controllo, ma si 

estende anche ai sistemi di raffreddamento secondario e richiede 

un raffreddamento spray più efficiente e affidabile. 

Un’attenta progettazione dei sistemi di raffreddamento secondario, 

che includa il posizionamento della zona di raffreddamento, 

la disposizione e selezione degli ugelli, come 

pure l’impiego di adeguati sistemi di controllo del raffreddamento 

secondario possono soddisfare queste necessità. 

La riduzione al minimo dei tempi morti per la manutenzione 

rappresenta un fattore chiave nella massimizzazione della 

produzione; l’utilizzo dei più innovativi sistemi di collettori 

e la disposizione degli ugelli possono contribuire a minimizzare 

la manutenzione dell’impianto di raffreddamento secondario. 

Questi sistemi di collettori prevedono un assemblaggio 

rigido e auto-allineante dell’ugello, che ne assicura 

sia l’allineamento che l’accessibilità ai fini della manutenzione. 

52 


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Development of high grade 

seamless pipes for deepwater application 

by metallurgical design 

E. Anelli, D. Colleluori, G. Cumino, A. Izquierdo, H. Quintanilla 

New solutions for the metallurgical design of high performance quenched and tempered (Q&T) seamless 

pipes of high grades from X65 up to X80 were found throughout a systematic work involving 

metallurgical modelling, laboratory tests, pilot and industrial trials. Both linepipes and risers for 

deepwater offshore fields such as Gulf of Mexico were considered. The target microstructure in the asquenched 

condition has been identified as refined low-C bainite/martensite matrix (> 70%). 

This is promoted through the control of austenite grain growth during the heating stage (austenitization 

before quenching), proper alloy additions and through a very effective quenching. Promising low-alloy 

steels and suitable quenching and tempering conditions identified by metallurgical modelling were 

verified by laboratory heats (80 kg ingots) that were processed at a pilot scale and submitted to 

microstructural examination and mechanical testing. The best solutions were used in preliminary 

industrial trials, also utilised for a fine tuning. The Ni-Cr-Mo-Nb-V alloy system showed very interesting 

combinations of strength-toughness and field weldability, suitable for the production of heavy wall X65 

grade linepipes for sour service and X80 top tension risers. 

Memorie 

Key words: Seamless pipe, linepipe, riser, strength, toughness, microstructure, quenching, tempering, metallurgical modelling 

INTRODUCTION 

The technological evolution in the offshore sector exhibits a 

trend towards an increasing use of high strength steels (grade 

X65 to X80, and higher) both for risers and flowlines, 

although the service conditions and the performance required 

for the two systems are different. This trend is supported 

both by economical and technical reasons, because the development 

of deepwater oil and gas reserves is continuously 

facing the challenge of containing/reducing costs in all components. 

[1-2] 

For instance, riser system costs are quite sensitive to water 

depth and there is a need to explore new technical solutions 

and reduce raiser weight for ultra-deep water environments 

(greater than 2000 m). The use of high strength steels can 

decrease the wall thickness up to 30%, resulting in a more 

efficient design. Thinner wall risers mean reduced buoyancy 

requirements and less hydrodynamic loading on these components, 

with consequent improvement in riser response. 

For large field developments employing floating production 

facilities, with many heavy risers attached directly to the 

surface structure, payload limitations will receive higher 

consideration. Therefore, in this context the availability of 

higher-grade weldable steel risers, with a wall thickness to 

outside diameter ratio (WT/OD) adequate to the expected 

collapse performance is of engineering importance. 

On the other hand, flowline wall thickness is increasing to 

provide sufficient resistance for the very high operating 

pressures. The trend in flowline specifications for deepwater 

E. Anelli, D. Colleluori - Centro Sviluppo Materiali S.p.A. - Rome, Italy 

G. Cumino - TenarisDalmine - Dalmine (BG), Italy 

A. Izquierdo, H. Quintanilla - TenarisTamsa - Veracruz, Mexico 

Paper presented at the 2 nd International Conference on New Developments 

in Metallurgical Process Technology, 

Riva del Garda, 19-21 September 2005, organised by AIM 

offshore fields is a consequence of both complex oil-gas 

field conditions, such as high pressure and high temperature 

(HPHT) and developments in design criteria (i.e. limit state 

design), welding and laying technologies. Often the requirements 

are close to the manufacturing limit of welded pipes, 

therefore seamless pipes that allow a higher (WT/OD) are 

preferred. 

As a matter of fact pipe manufacturers are facing new challenges 

coming from new and/or more demanding material 

requirements, often related to specific performances and applications, 

including sour service which set limitations such 

as maximum material hardness (HV ≤ 248). 

During the past years technologies have been developed in 

the field of quenched and tempered (Q&T) seamless pipe. In 

particular, the heat treatment capabilities of heavy wall pipes 

have been improved through the introduction of external 

and internal water quenching, which decreases the throughthickness 

temperature gradient. 

Modern seamless pipes can combine high strength (grade 

X65) with good toughness properties and good girth weldability. 

For instance, in the case of pipes with WT = 15 to 34 

mm, a reasonable balance between customer specifications, 

processing capabilities and product properties was found for 

low-C low alloy steels with 1%Mn and optimized contents 

of Mo, Nb and V. Such Q&T seamless pipes, which also 

showed good resistance to strain aging and HIC, were successfully 

delivered for HPHT offshore production lines. [3] 

However, for pipe wall thickness greater than 34 mm, the required 

strength (grade X65) cannot be easily achieved maintaining 

the required toughness level. Similar difficulties are 

experienced in the case of WT = 15-25 mm for higher 

strength levels (e.g. grade X80). Therefore, new solutions 

which are outside of the conventional pattern for (micro)-alloying 

additions followed so far for Q&T seamless pipes, 

have to be found for high performance seamless pipes throughout 

a more systematic work. 

In this paper, a description of the results of studies on high 

strength steel materials manufactured by Q&T processing is 

given. This work represents an on going development pro- 

1/2005 


PROCESSI 


gram on high performance Q&T seamless pipes for special 

deep water applications, involving metallurgical modeling, 

laboratory tests, pilot and industrial trials. The role of chemical 

composition and Q&T process conditions on microstructure 

and precipitation has been investigated, together 

with their effect on strength and toughness. These results have 

been exploited for the production of sour service grade 

API 5L X65/X70 for heavy wall flowlines (WT from 30 up 

to 42 mm) and X80 risers (WT = 16 to 25 mm) for deepwater 

offshore fields. 

METALLURGICAL BACKGROUND AND MODELLING 

Seamless pipes of medium O.D., i.e. up to 406 mm (16”) are 

presently produced by a hot rolling process carried out in the 

following main stages: hot piercing, rolling at retained mandrel 

mill and sizing. Quenching and tempering treatments 

are performed on the pipes in order to refine the microstructure 

and obtain the required properties. [3,4] 

A rational approach to the design and production of these 

materials requires the quantitative knowledge of the effects 

of steel chemistry and heat treatment variables on the microstructure 

and final mechanical properties. The influence of 

microalloying additions and Q&T practices on austenite refinement, 

phase transformation and response to heat treatments 

of low-C steels for seamless linepipes was investigated 

by dilatometry and pilot trials. 

Also an integrated model, containing a thermal routine for 

simulating pipe quenching, based on the integration by finite 

differences of the general Fourier heat equation, coupled 

with a microstructural model, has been applied for the design 

of both the chemical composition and Q&T conditions of 

seamless pipes. The thermal-metallurgical model is able to 

calculate the fraction of microstructural constituents and 

hardness of a steel subjected to rapid continuous cooling after 

austenitisation (i.e. quenching). The calculation is carried 

out by an Artificial Neural Network (ANN) trained on a selected 

database of CCT diagrams of linepipe steels. [5,6] 

ANN is a powerful tool to obtain empirical non-linear models 

of complex phenomena whose analysis by ordinary statistical 

techniques or by modeling through fundamental physical 

process is not possible or very difficult. [7] 

The program is able also to simulate a subsequent tempering 

treatment, predicting hardness (HV), yield strength (YS) 

and ultimate tensile strength (UTS) by an empirical approach. 

Different modules, each one describing an elementary 

process (e.g. austenitizing, quenching, tempering) can be 

managed by a user-friendly interface which allows to select 

the steel chemical composition, pipe diameter and wall 

thickness, and to set-up the process conditions. [5] 

Austenitizing 

The austenite grain size (AGS) depends on the austenitizing 

temperature and holding time, nature and size distribution of 

precipitates present in the as-rolled pipe. The more uniform 

the as-rolled microstructure, the easier it would be to homogenize 

the austenite. 

Laboratory tests and industrial trials have shown that to 

avoid the formation of coarse austenite grains (AGS > 25 

µm) in low carbon steels (0.08-0.11%C) the heating temperature 

has to be lower than 900 °C for C-Mn steels; however, 

V-Nb steels and V-Ti steels can be safely austenitized up 

to 920 to 950 °C to dissolve V-rich precipitates, without problems 

due to the pinning effect of Nb (C, N) and Ti (C, N) 

fine particles which hinder grain boundary movement. 

Quenching 

Of concern in a quenching process of seamless pipes are the 

effects of through-thickness cooling rate gradients, induced 

by surface water cooling, and the sequence of transformation 

and the resultant microstructure and hardness profile. A 

specific test program was performed with the main objective 

of measuring the phase transformation characteristics of austenite 

under continuous cooling conditions by dilatometry 

and metallography (construction of CCT diagrams). Mathematical 

modeling, which links the basic principles of heat 

transfer and microstructural phenomena was effectively applied 

in this field. 

The volume fraction of microstructural constituents and 

hardness of as-quenched linepipes were predicted as a function 

of the local cooling rate, calculated by the thermal model, 

by means of ANN. 

Input data of ANN are the chemical composition of the 

steel, the austenitizing temperature, the austenite grain size 

and the cooling rate (CR). The output contains information 

on the as-quenched state of the steel in terms of harness and 

amount of microstructural constituents (e.g. ferrite, pearlite, 

bainite, martensite). The ANN was trained on a huge database 

of continuous cooling diagrams of Nb-V micro-alloyed 

linepipe steels (C=0.05-0.2%, Mn=0.5-2.0%), with possible 

Mo, Ni, and Cr alloy additions. Standard deviations of the 

regression lines between computed and experimental fractions 

were generally below 5%. The standard deviation for 

the hardness was 11 Vickers. In all cases the associated correlation 

coefficient was greater than 0.91. [5,6] 

In order to check the prediction capability of ANN, new experimental 

CCT diagrams, not used for training, were determined 

on a Nb-V-microalloyed pipeline steel. Two diagrams 

were determined by austenitising the steel for 10 minutes at 

930 and 1100 °C in order to obtain AGS of 9 and 40 mm, respectively. 

Hardness of the as-quenched samples were measured 

and the related microstructures analysed by optical 

microscope to evaluate the volume fractions of the microstructural 

constituents. Experimental data were compared to 

the predictions of the microstructural ANN model. The good 

results are presented in Figures 1 and 2. 

Other experiments, including quenching of pipes by various 

industrial devices, were carried out. Both pipes instrumented 

by thermocouples and through-thickness hardness profiles 

on as-quenched pipes were employed to tune the heat 

transfer coefficient as a function of process conditions (e.g. 

water flow, etc.). 

Fig. 1 – Comparison between the ANN predictions and the 

experimental data for hardness in a Nb-V linepipe steel. 

Fig. 1 – Confronto tra previsioni con la rete artificiale neuronale e 

dati sperimentali sulla durezza di acciai al Nb-V per linepipe. 

34 


PROCESSI 

Fig. 2 – Comparison between 

the ANN predictions and the 

experimental data for ferrite 

and bainite volume fractions 

in a Nb-V linepipe steel. 

Fig. 2 – Confronto tra 

previsioni con la rete 

artificiale neuronale e dati 

sperimentali sulla percentuale 

di ferrite e bainite in acciai al 

Nb-V per linepipe. 

Memorie 

Simulations carried out by the model and industrial trials 

showed that linepipes with wall thickness less than 16 mm 

can be effectively quenched passing continuously through 

water jets produced by nozzles arranged in a series of rings 

forming a tunnel. In the case of heavy wall pipes, external 

and internal quenching is needed to reduce hardness gradients 

and make more homogeneous the as-quenched structure. 

This type of quenching is carried out by dipping the pipe 

in a tank containing stirred water. During quenching the 

pipe is under rotation and an internal water jet is used, too. 

Tempering 

The tempering conditions, mainly temperature and the presence 

of elements able to give precipitation hardening and to 

slow the recovery/recrystallization process, are the controlling 

factors for the combination of strength and toughness, 

for a given as-quenched microstructure. 

The integrated model system is also able to simulate a subsequent 

tempering treatment of the microstructure after 

quenching. In this case an empirical approach was used, 

which permits the estimation of the final hardness taking into 

account also the effect of secondary hardening phenomena 

due to precipitation of second phases in steels containing 

V and/or Mo. The yield strength (YS) and the ultimate tensile 

strength (UTS) of the Q&T material were estimated from 

hardness using empirical equations. [6] 

A comparison of calculated and experimental strength values 

of as-quenched industrial specimens submitted to tempering 

under very well controlled conditions in the laboratory 

are shown in Table I. 

Sensitivity Analysis 

A sensitivity analysis was performed on a reference steel, 

Fig. 3 – Calculated effect of AGS and cooling rate during 

quenching on hardness. 

Fig. 3 – Effetto della dimensione media del grano austenitico e 

della velocità di raffreddamento durante tempra sulla durezza. 

Valori calcolati da modello. 

having the chemical composition shown in Table II, in order 

to identify the role of chemical composition and process 

conditions on the phase transformation and response to tempering. 

The well known effect of the austenite grain size in enhancing 

the hardenability of the steel clearly appears from Fig. 3. 

Hardness increases monotonically with AGS and with coo- 

Table I – Comparison between predicted and 

experimental strength for a Nb-V linepipe steel 

submitted to tempering at various temperatures 

for 60 min after industrial quenching (WT = 22 

mm). 

Tabella I – Confronto tra resistenza meccanica 

misurata e quella calcolata per tubi in acciai al 

Nb-V per linepipe rinvenuti a varie temperature 

per 60 min dopo tempra industriale. (spessore 

22 mm). 

Table II –Chemical composition (mass %) of 

the reference steel used in the sensitivity 

analysis. 

Tabella II – Analisi chimica (% in massa) 

dell’acciaio di riferimento usato per l’analisi di 

sensibilità. 

Tempering HV YS (MPa) UTS (MPa) 

Temperature (°C) Calc Exp Calc Exp Calc Exp 

620 229 229 617 605 695 706 

640 219 225 585 576 670 681 

660 210 221 558 586 647 665 

680 202 217 530 561 624 637 

700 194 209 503 516 603 615 

CMn Si Ni Cr Mo P S Cu V Al Nb 

0.12 1.2 0.3 0.1 0.15 0.07 0.01 0.004 0.15 0.05 0.025 0.02 

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Fig. 4 – Calculated effect of AGS and cooling rate during 

quenching on microstructural constituents. 

Fig. 4 – Effetto della dimensione media del grano austenitico e 

della velocità di raffreddamento durante tempra sulle frazioni dei 

costituenti microstrutturali. Valori calcolati da modello. 

Fig. 5 – Calculated effect of carbon content and cooling rate 

during quenching on hardness. 

Fig. 5 – Effetto del tenore di carbonio e della velocità di 

raffreddamento durante tempra sulla durezza. Valori calcolati da 

modello. 

Fig. 6 – Calculated effect of manganese content and cooling rate 

during quenching on hardness. 

Fig. 6 – Effetto del tenore di manganese e della velocità di 

raffreddamento durante tempra sulla durezza. Valori calcolati da 

modello. 

Fig. 7 – Calculated effect of manganese content and AGS during 

cooling at 10 °C/s on microstructural constituents. 

Fig. 7 – Effetto del tenore di manganese e della dimensione media 

del grano austenitico sulle frazioni dei costituenti microstrutturali 

formatisi per velocità di raffreddamento di 10 °C/s. Valori 

calcolati da modello. 

Fig. 8 – Calculated effect of ferrite amount on YS of Q&T 

materials (tempering at 660 °C for 60 min). 

Fig. 8 – Effetto della percentuale di ferrite sulla tensione di 

snervamento di materiali temprati e rinvenuti a 660 °C per 60 

min. Valori calcolati da modello. 

ling rate, showing a saturation characterized by a stabilization 

in the amounts of bainite and martensite in the as-quenched 

microstructure with coarse grain (Fig.4). At the same 

time, as expected, ferrite appears to be present at low cooling 

rates and small AGS and to decrease as both cooling rate 

and AGS increase. One can observe that the bainite fraction 

goes through a maximum due to competition with ferrite 

at small AGS and martensite in the large grain region 

(Fig.4). 

It can be noticed that the ANN has been extrapolated in the 

range of grain sizes from 50 to 100 mm with respect to the 

training interval for AGS. Apparently this has not caused 

any saturation effect in the output parameters. 

Also the increase of carbon content (Fig.5) has shown a similar 

effect on hardness as that exhibited by AGS, in agreement, 

also in this case, with the experience. 

In the third example the effect of Mn is considered for different 

values of AGS and cooling rate (Fig.6). The hardening 

effect of Mn is associated with the promotion of bainitic 

structures and this behavior is favored by large grains 

(Fig.7). 

The application of the mathematical model indicates that to 

attain the required yield strength for grade X65 in the case 

of 40 mm (i.e. CR of 12-15 °C/s) it is necessary to have a 

fraction of polygonal ferrite below 30% (Fig.8). 

The model is a very powerful and easy-to-use tool for designing 

the industrial Q&T processes taking into account all 

36 


PROCESSI 

the relevant process parameters. It has shown a very good 

reliability in reproducing the experimental data not only on 

the microstructure, but also on mechanical properties. For 

Q&T seamless linepipes, it is able to quantify both the through-thickness 

thermal gradients during the quenching 

process together with the related local microstructures and 

hardness, and the effect of subsequent tempering treatments. 

The model is an effective tool in defining the optimum Q&T 

treatments for a given steel to match the required tensile 

properties of the final product. However, no information on 

toughness is available from modelling. Therefore, a specific 

experimental activity has been designed and carried out to 

assess the effect of microstructure and precipitation on 

strength-toughness combination, starting from a promising 

chemical composition identified by metallurgical modelling. 

Memorie 

HEAVY WALL HIGH STRENGTH FLOWLINES 

Pilot Trials 

A series of laboratory heats, with designed changes in the 

content of Mo, Ni, Cr, V and C, with respect to a base steel 

composition, were vacuum cast as 80 kg ingots. The carbon 

equivalent, Ceq (IIW) was in the range 0.33% to 0.43% and 

maximum Pcm parameter was 0.25%. 

The ingots were hot rolled by a pilot mill simulating the typical 

thermo-mechanical process of heavy wall seamless pipes 

(40 mm final thickness). All ingots were instrumented 

with thermocouples and the evolution of temperature during 

hot rolling was recorded. 

The hot rolled materials were quenched in stirred water and 

tempered under strictly controlled parameters (Table III), 

being each piece instrumented by a thermocouple embedded 

at mid-thickness. 

Fig. 9 – Strength vs polygonal ferrite amount for various 

laboratory Q&T steels. 

Fig. 9 – Resistenza meccanica in funzione della percentuale di 

ferrite poligonale per vari acciai di laboratorio. 

Min 

Max 

Austenitizing Temperature (°C) 920 1020 

Cooling rate during quenching (°C/s) 10 25 

Tempering Temperature (°C) 630 690 

Table III – Range of laboratory heat treatment conditions. 

Tabella III – Intervallo esplorato per le condizioni di trattamento 

termico in laboratorio. 

The Q&T materials were examined by light and scanning 

electron microscopy. Microstructures were observed on sections 

after 2%-nital etching. Islands of high carbon martensite 

with retained austenite (MA constituent) were revealed 

by selective etching. [8] 

The austenite grain boundaries were revealed by etching in a 

saturated aqueous picric acid solution containing a few drops 

of teepol and HCl. 

The average austenite grain size (AGS) was measured according 

with ASTM E112. 

Tensile and Charpy V-notch testing was conducted on transverse 

specimens. Charpy-V transition curves were determined 

together with the fracture appearance transition temperature 

(50% FATT). 

Effect of Alloy Design on Microstructure 

It was confirmed, as suggested by model simulations that in 

order to achieve the required yield strength level after tempering 

it is a pre-requisite to maintain the fraction of polygonal 

ferrite well below 30% in the as-quenched material 

(Fig.9). 

A judicious addition of Mo, Ni and Cr allows to develop after 

quenching a microstructure containing a suitable combination 

of constituents such as fine bainite (B > 75%), poly- 

Fig. 10 – Mo-Ni-Nb-V Steel. AGS obtained with austenitizing 

temperature of 920°C and quenching with CR = 12°C/s. Prior 

austenite grain boundaries were revealed by selective etching by 

saturated picric solution. 

Fig. 10 – Acciaio al Mo-Ni-Nb-V. Grano austenitico ottenuto dopo 

trattamento a 920 °C e tempra con velocità pari a 12°C/s. I bordi 

dei grani austenitici primitivi sono stati evidenziati mediante 

attacco metallografico selettivo mediante soluzione acquosa 

satura di acido picrico. 

gonal ferrite (PF < 25%) and fine islands of MA constituent, 

uniformly dispersed in the matrix. This predominantly bainitic 

structure was found to exhibit good toughness values 

especially when the AGS was fine (< 15 mm; > ASTM 

No.9) and homogeneous. 

An example of the typical AGS (ASTM No.9.2) of Mo-Ni- 

Nb-V steel, quenched from 920 °C, is reported in Fig.10. 

The addition of Nb slows down grain growth and helps to 

maintain relatively fine and homogeneous the AGS during 

austenitization. 

In the case of fine AGS, the increase of the cooling rate from 

12 °C/s to 20 °C/s, refines the microstructure (Fig.11). 

Therefore, even better strength-toughness combinations are 

expected for linepipes with thickness smaller than 40 mm. 

Concerning the microstructure evolution during tempering, 

a better spheroidisation of cementite and a more extended 

transformation of MA islands into ferrite and carbides, was 

observed with increasing tempering temperature (Fig.12). 

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Fig. 11 – Mo-Ni-Nb-V Steel. Microstructure obtained after austenitizing at 920°C and quenching with CR=12°C/s and 20 °C/S, respectively. 

The as-quenched microstructure was revealed by 2% Nital etching. a) CR = 12 °C/s b) CR = 20 °C/s. 

Fig. 11 – Acciaio al Mo-Ni-Nb-V. Microstruttura ottenuta dopo austenitizzazione a 920°C e tempra con velocità di raffreddamento di 12°C/s 

e 20 °C/S, rispettivamente. La microstruttura grezza di tempra è stata evidenziata con attacco al nital 2%. 

a 

Fig. 12 – Cementite morphology after Q&T (SEM image). a) Low Tempering Temperature b) High Tempering Temperature. 

Fig. 12 – Morfologia della cementite dopo tempra e rinvenimento (immagini al microscopio elettronico a scansione). 

Strength and Toughness Properties 

The general pattern of strength/toughness properties as a 

function of (micro)-alloy design is reported in Fig.13. The 

Mo-Ni-Nb-V composition gives excellent strength/toughness 

combination: yield strength well above 460 MPa (grade 

X65); 50%FATT as low as – 85°C. 

With reference to the same base composition and Q&T conditions 

of Fig.13: 

• The increase of C content up to 0.13% gives strengthening 

(DYS = + 60 MPa), but is detrimental to toughness and 

weldability. 

• The V-free version (Mo-Ni-Nb steel) allows to reduce 

FATT, but at expense of strength (DYS = – 35 MPa). 

• Addition of Cr produces further improvement in toughness 

(50% FATT below – 100°C). 

Therefore the Mo-Ni-Cr-Nb-V-steel resulted to be the most 

promising for the production of heavy wall linepipes. 

The laboratory materials exhibit the general trend of increasing 

yield to tensile (Y/T) ratio when yield strength increases. 

For the required strength levels (YS > 450 MPa) the 

Y/T values are in the range 0.83 to 0.87, being the highest 

ratios related to a predominantly bainitic structure. 

The alloy design based on the Mo-Ni-Cr-Nb-V version exhi- 

b 

Fig. 13 – Toughness vs yield strength for various laboratory Q&T 

steels. 

Fig. 13 – Tenacità in relazione alla tensione di snervamento per 

vari acciai di laboratorio temprati e rinvenuti. 

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PROCESSI 

bits the best results in terms of low values of Y/T ratio. 

Concerning toughness, when the as-quenched microstructure 

is fully bainitic and the volume fraction of MA is very 

low, the 50% FATT of Q&T material is practically independent 

of the tempering temperature, as expected. The increase 

of tempering temperature is effective in improving toughness 

if significant amounts of MA constituent, in the form of 

large islands, are present in the as-quench material. In this 

case, usually, the rising of tempering temperature from 

630°C to 680°C leads to a decrease in yield strength (DYS = 

– 20 MPa) and a slight improvement of toughness in terms 

of 50% FATT (DFATT = – 10°C). 

Industrial Production and Pipe Qualification 

Heavy wall seamless pipes of medium diameter (OD = 219 

to 323 mm) and WT = 30 to 40 mm were produced at Tenaris 

works by the seamless process, using the steel chemistry 

range and heat treatment conditions identified as promising 

by the metallurgical design. Water quenching was carried 

out by dipping the rotating pipe in a tank containing stirred 

water. Also an internal water jet is used to increase heat transfer 

at the inner surface. 

All pipes were manufactured according to specific customer 

requirements for production risers and flowlines. In particular, 

in addition to weldability requirements, in terms of carbon 

equivalent, Ceq (IIW) and Pcm parameter, suitable tensile 

properties at room and at 130 °C shall be guaranteed 

(minimum yield strength = 448 MPa). 

The characterization of selected pipes from the production 

was carried out by extensive metallography and mechanical 

testing, which included hardness measurements, longitudinal 

and transverse tensile testing, Charpy-V impact testing, 

crack tip opening displacement (CTOD) testing. 

The industrial production confirmed the effect of main process 

parameters and metallurgical factors on microstructure 

and strength-toughness combination, as outlined by the laboratory 

experiments, and allowed to identify the actions for 

a fine tuning to develop a good combination of strength and 

toughness. 

All hardness values on Q&T pipes were below 248 HV10. 

The mechanical properties of various industrial materials in 

terms of 50% FATT and yield strength are summarized in 

Fig.14. 

The materials produced using the Mo-Ni-Cr-Nb-V steel, have 

significantly improved toughness for a given yield 

strength level between 460 and 530 MPa, compared to the 

conventional chemistry without Ni. The productions indicate 

that suitable toughness levels (i.e. 50% FATT < – 50°C) 

have been achieved. Also high CTOD values at – 10 °C (> 

1.1 mm) were measured on the Q&T seamless pipes. Best 

results were attributed to the following aspects: 

• strict control of process parameters during heating in order 

to develop uniform and small PAGS; 

• effective quenching to promote higher volume fractions of 

bainite and a more refined final microstructure. 

Suitable strength levels of pipes for flowlines and produc- 

Fig. 14 – Toughness vs yield strength for heavy wall seamless 

pipes. 

Fig. 14 – Tenacità in relazione alla tensione di snervamento per 

tubi senza saldatura di grosso spessore, per linepipe. 

tion risers were also achieved at 130 °C. 

Some pipes were girth welded by GMAW using low (0.6 

kJ/mm) and high (3 kJ/mm) heat inputs. CTOD was also 

performed in the HAZ of girth welds performed by GMAW. 

CTOD specimens were located at both the Transformed- 

HAZ (THAZ) and Visible-HAZ boundary (VHAZ). Post-test 

validity checks were carried out by specimen sectioning 

followed by fractographic and metallographic evaluation. 

The CTOD results at 4 °C were quite good both for the 

THAZ (CGHAZ), with values of 0.6 - 1 mm and the VHAZ 

with values of 1 - 1.5 mm. At – 10 °C, CTOD remained 

always greater than 0.3 mm. 

HIGH STRENGTH RISERS (WT = 15 to 25 mm) 

Due to the higher cooling rates during quenching, these pipes 

can develop a predominantly bainitic-martensitic structure 

more easily. The tuning of Mn, Mo, Cr, V and Ni additions 

was done on the basis of the results from the metallurgical 

model and pilot trials (Table I). 

A steel with carbon equivalent, Ceq° (IIW) 

, in the range 

0.37% to 0.40% and maximum Pcm parameter below 0.2% 

was selected. 

The sizes and mechanical properties of the risers produced 

are shown in Table IV. 

Hardness near the internal surface is usually higher because 

a high-pressure water jet blows the steam generated inside 

the pipe during the immersion in the water tank, promoting 

a higher heat transfer coefficient, thus increasing the severity 

of quenching. However, it is important to mention that 

suitable hardness can be obtained adjusting the tempering 

Memorie 

Longitudinal * Transverse ** 

Item OD WT Ceq° (IIW) 

HV YS UTS El FAT YS UTS El FAT 

(mm) (mm) max (MPa) (MPa) (%) T (°C) (MPa) (MPa) (%) T (°C) 

1 323.9 16 0.38 243 589 677 42 – 60 596 672 25.4 – 50 

2 298.5 22 0.40 248 594 678 41 – 50 602 681 25.0 – 40 

* full strip specimen; ** round specimen 

Table IV – Sizes and average mechanical properties of the produced risers. 

Tabella IV – Dimensioni dei riser prodotti e relative proprietà meccaniche. 

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treatment parameters. As all the hardness values were below 

250 HV10, these risers can be used for sour service applications. 

An interesting feature found in these pipes is that there is not 

a significant difference between the longitudinal and the 

transverse tensile properties at room temperature. This is 

due to the high isotropy which is typical of seamless pipes. 

The materials exhibited an absorbed energy value above 140 

J up to –60 °C. This is a consequence of the microstructure 

promoted after external/internal quenching and tempering. 

Also high CTOD values (> 1.2 mm) were measured at – 20 

°C on these high strength Q&T seamless pipes. 

Weldability trials proved that the selected chemical composition 

combined with proper welding procedure assures 

hardness values lower than 280 HV and satisfactory toughness 

levels both in the weld metal and heat affect zone. 

CONCLUSIONS 

New chemistries and optimized Q&T conditions were identified 

for the production of seamless pipes for both flowlines 

with thickness up to 42 mm and risers of grade X80 and 15 

to 25 mm WT, through an extensive characterization of laboratory 

and industrial materials. A number of selected alloy 

systems have been systematically investigated. Main conclusions 

are: 

• The as-quenched microstructure plays a primary role. The 

best toughness values are related to a predominantly bainitic 

microstructure after quenching (bainite > 70%) combined 

with a homogeneous and fine distribution of MA constituent. 

This is promoted through the control of austenite 

grain growth during the heating stage and an effective 

quenching. 

SVILUPPO METALLURGICO DI TUBI SENZA SALDATURA 

DI GRADO ELEVATO PER APPLICAZIONI IN ACQUE PROFONDE 

PAROLE CHIAVE: Tubi senza saldatura, linepipe, riser, 

resistenza meccanica, tenacità, microstruttura, tempra, 

rinvenimento, modelli metallurgici 

Attraverso un’indagine sistematica condotta impiegando sia la 

modellistica metallurgica sia sperimentazioni di laboratorio e 

prove su impianto pilota e in scala industriale, sono state identificate 

nuove soluzioni per la progettazione metallurgica di tubi 

senza saldatura temprati e rinvenuti (Q&T) ad elevate prestazioni 

ed alto grado (da X65 a X80). Un approccio razionale alla 

progettazione e alla produzione di questi materiali ha richiesto la 

conoscenza quantitativa degli effetti dell’analisi chimica dell’acciaio 

e dei parametri di trattamento termico sulla microstruttura 

e sulle proprietà meccaniche finali. L’influenza degli elementi di 

lega e delle condizioni di tempra e di rinvenimento sull’affinamento 

dell’austenite, sulle trasformazioni di fase e sulla risposta 

ai trattamenti termici è stata studiata mediante apposite sperimentazioni 

su acciai a basso tenore di carbonio impiegati nella 

fabbricazione di tubi senza saldatura per linepipe e riser. Inoltre, 

è stato validato un modello matematico integrato, costituito da 

un modulo termico per la simulazione numerica della tempra in 

acqua dei tubi basato sulla soluzione dell’equazione di Fourier 

del trasporto di calore e da un modello microstrutturale. Questo 

è stato applicato alla progettazione sia della composizione dell’acciaio 

sia del trattamento termico dei tubi senza saldatura. Il 

modello termico-metallurgico è in grado di calcolare la frazione 

dei costituenti microstrutturali e la durezza di tubi in acciai sottoposti 

ad un raffreddamento continuo rapido dopo austenitizzazione 

(tempra). Il calcolo della trasformazione di fase viene effettuato 

mediante una rete neuronale artificiale (ANN), addestrata 

ABSTRACT 

• The Mo-Ni-Cr-Nb-V alloy system showed the best combination 

of strength-toughness and weldability. 

• The tempering temperature has a secondary role. However, 

higher tempering temperature leads always to a slightly 

lower yield strength and, in the case of coarse MA 

islands and cementite particles, a slightly improved toughness. 

• The CTOD results in the HAZ of GMAW girth joints were 

good. 

REFERENCES 

1. M. Cook, OPT USA 2002, Houston, 16-17 Sept. 2002. 

2. U. Eigbe, OPT USA 2002, Houston, 16-17 Sept. 2002.4. 

3. E. Anelli et al., 11th International Offshore and Polar Engineering 

Conference and 2nd International Deep Ocean 

Technology Symposium, Stavanger, Norway, June 17-22, 

2001. 

4. J.C. Gonzalez, M. Tivelli, H. Quintanilla, G. Cumino, 

Proc. 40th Mechanical Working and Steel Processing 

Conference, ISS, 1998, pp 621-626. 

5. P.E. Di Nunzio, E. Anelli and G. Cumino, EUROMAT 

2001 Int. Conf., Rimini, Italy, 10-14 June 2001, paper no. 

546. 

6. E. Anelli, M.C. Cesile, P.E. Di Nunzio, Int. Conf. “Modeling 

Control and optimization in Ferrous and non Ferrous 

industry”, ed. F. Kongoli, B.G. Thomas, K. Sawamiphakdi, 

ISS & TMS, Chicago, 9-12 Nov. 2003, pp. 453-466. 

7. Special Issue “Application of Neural Network Analysis in 

Materials Science”, ISIJ Int., Vol.39, No.10, 1999, pp. 

965-1105. 

8. D. Colleluori, Pract. Met., Vol.20, 1983, pp. 546-553. 

su un database costituito da una selezione di diagrammi di trasformazione 

di fase in raffreddamento continuo (CCT) di acciai 

per linepipe. Il programma consente di simulare anche il successivo 

trattamento di rinvenimento, prevedendo, mediante un approccio 

empirico, durezza (HV), tensione di snervamento (YS) e 

carico unitario a rottura (UTS). Vari moduli, ognuno in grado di 

descrivere un processo elementare (ad es. austenitizzazione, tempra, 

rinvenimento), sono gestiti da un’interfaccia-utente di uso 

immediato che consente di selezionare l’analisi chimica, il diametro 

e lo spessore del tubo e di stabilire le condizioni di trattamento 

termico. L’applicazione della modellistica matematica ha 

mostrato che per ottenere la tensione di snervamento desiderata 

per il grado X65, nel caso di tubi di spessore 40 mm, è necessario 

sviluppare una percentuale di ferrite poligonale inferiore al 

30% in volume. La microstruttura ottimale dopo tempra, costituita 

prevalentemente da bainite-martensite a basso carbonio, 

viene promossa da un controllo delle dimensioni dei grani austenitici 

durante l’austenitizzazione che precede la tempra, da aggiunte 

calibrate di elementi di lega e da un raffreddamento molto 

efficace. Attraverso l’applicazione del modello metallurgico sono 

state definite alcune composizioni promettenti di acciai bassolegati 

e le condizioni di tempra e rinvenimento più idonee. Le 

previsioni sono state verificate con colate di laboratorio (lingotti 

da 80 kg) processate su scala pilota, esaminate metallograficamente 

e valutate in termini di proprietà meccaniche con particolare 

attenzione alla tenacità. I materiali più promettenti sono stati 

impiegati per prove su scala industriale e per la calibrazione 

fine dei modelli matematici. Acciai al Ni-Cr-Mo-Nb-V hanno 

mostrato combinazioni di resistenza meccanica e tenacità molto 

interessanti per la produzione di tubi di grado API 5L X65/X70 

per flowline di grosso spessore (da 30 a 40 mm) e riser (di spessore 

da 16 a 25 mm) di grado fino a X80, per lo sfruttamento di 

giacimenti in acque profonde. 

40

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